voltage of the ball mill 75 kw

grinding circuit - an overview | sciencedirect topics

grinding circuit - an overview | sciencedirect topics

Grinding circuits are fed at a controlled rate from the stockpile or bins holding the crusher plant product. There may be a number of grinding circuits in parallel, each circuit taking a definite fraction of the feed. An example is the Highland Valley Cu/Mo plant with five parallel grinding lines (Chapter 12). Parallel mill circuits increase circuit flexibility, since individual units can be shut down or the feed rate can be changed, with a manageable effect on production. Fewer mills are, however, easier to control and capital and installation costs are lower, so the number of mills must be decided at the design stage.

The high unit capacity SAG mill/ball mill circuit is dominant today and has contributed toward substantial savings in capital and operating costs, which has in turn made many low-grade, high-tonnage operations such as copper and gold ores feasible. Future circuits may see increasing use of high pressure grinding rolls (Rosas et al., 2012).

Autogenous grinding or semi-autogenous grinding mills can be operated in open or closed circuit. However, even in open circuit, a coarse classifier such as a trommel attached to the mill, or a vibrating screen can be used. The oversize material is recycled either externally or internally. In internal recycling, the coarse material is conveyed by a reverse spiral or water jet back down the center of the trommel into the mill. External recycling can be continuous, achieved by conveyor belt, or is batch where the material is stockpiled and periodically fed back into the mill by front-end loader.

In Figure 7.35 shows the SAG mill closed with a crusher (recycle or pebble crusher). In SAG mill operation, the grinding rate passes through a minimum at a critical size (Chapter 5), which represents material too large to be broken by the steel grinding media, but has a low self-breakage rate. If the critical size material, typically 2550mm, is accumulated the mill energy efficiency will deteriorate, and the mill feed rate decreases. As a solution, additional large holes, or pebble ports (e.g., 40100mm), are cut into the mill grate, allowing coarse material to exit the mill. The crusher in closed circuit is then used to reduce the size of the critical size material and return it to the mill. As the pebble ports also allow steel balls to exit, a steel removal system (such as a guard magnet, Chapters 2 and 13Chapter 2Chapter 13) must be installed to prevent them from entering the crusher. (Because of this requirement, closing a SAG mill with a crusher is not used in magnetic iron ore grinding circuits.) This circuit configuration is common as it usually produces a significant increase in throughput and energy efficiency due to the removal of the critical size material.

An example SABC-A circuit is the Cadia Hill Gold Mine, New South Wales, Australia (Dunne et al., 2001). The project economics study indicated a single grinding line. The circuit comprises a SAG mill, 12m diameter by 6.1m length (belly inside liners, the effective grinding volume), two pebble crushers, and two ball mills in parallel closed with cyclones. The SAG mill is fitted with a 20MW gearless drive motor with bi-directional rotational capacity. (Reversing direction evens out wear on liners with symmetrical profile and prolongs operating time.) The SAG mill was designed to treat 2,065t h1 of ore at a ball charge of 8% volume, total filling of 25% volume, and an operating mill speed of 74% of critical. The mill is fitted with 80mm grates with total grate open area of 7.66m2 (Hart et al., 2001). A 4.5m diameter by 5.2m long trommel screens the discharge product at a cut size of ca. 12mm. Material less than 12mm falls into a cyclone feed sump, where it is combined with discharge from the ball mills. Oversize pebbles from the trommel are conveyed to a surge bin of 735t capacity, adjacent to the pebble crushers. Two cone crushers with a closed side set of 1216mm are used to crush the pebbles with a designed product P80 of 12mm and an expected total recycle pebble rate of 725t h1. The crushed pebbles fall directly onto the SAG mill feed belt and return to the SAG mill.

SAG mill product feeds two parallel ball mills of 6.6m11.1m (internal diameterlength), each with a 9.7MW twin pinion drive. The ball mills are operated at a ball charge volume of 3032% and 78.5% critical speed. The SAG mill trommel undersize is combined with the ball mills discharge and pumped to two parallel packs (clusters) of twelve 660mm diameter cyclones. The cyclone underflow from each line reports to a ball mill, while the cyclone overflow is directed to the flotation circuit. The designed ball milling circuit product is 80% passing 150m.

Several large tonnage copper porphyry plants in Chile use an open-circuit SAG configuration where the pebble crusher product is directed to the ball mills (SABC-B circuit). The original grinding circuit at Los Bronces is an example: the pebbles generated in the two SAG mills are crushed in a satellite pebble crushing plant, and then are conveyed to the three ball mills (Mogla and Grunwald, 2008).

Hydrocyclones have come to dominate classification when dealing with fine particle sizes in closed grinding circuits (<200m). However, recent developments in screen technology (Chapter 8) have renewed interest in using screens in grinding circuits. Screens separate on the basis of size and are not directly influenced by the density spread in the feed minerals. This can be an advantage. Screens also do not have a bypass fraction, and as Example 9.2 has shown, bypass can be quite large (over 30% in that case). Figure 9.8 shows an example of the difference in partition curve for cyclones and screens. The data is from the El Brocal concentrator in Peru with evaluations before and after the hydrocyclones were replaced with a Derrick Stack Sizer (see Chapter 8) in the grinding circuit (Dndar et al., 2014). Consistent with expectation, compared to the cyclone the screen had a sharper separation (slope of curve is higher) and little bypass. An increase in grinding circuit capacity was reported due to higher breakage rates after implementing the screen. This was attributed to the elimination of the bypass, reducing the amount of fine material sent back to the grinding mills which tends to cushion particleparticle impacts.

Circulation of material occurs in several parts of a mineral processing flowsheet, in grinding and flotation circuits, for example, as well as the crushing stage. In the present context, the circulating load (C) is the mass of coarse material returned from the screen to the crusher relative to the circuit final product (or fresh feed to the circuit), often quoted as a percentage. Figure 8.2 shows two closed circuit arrangements. Circuit (a) was considered in Chapter 6 (Example 6.1), and circuit (b) is an alternative. The symbols have the same meaning as before. The relationship of circulating load to screen efficiency for circuit (a) was derived in Example 6.1, namely (where all factors are as fractions):

The circulating load as a function of screen efficiency for the two circuits is shown in Figure 8.3. The circulating load increases with decreasing screen efficiency and as crusher product coarsens (f or r decreases), which is related to the crusher set (specifically the closed side setting, c.s.s.). For circuit (a) C also increases as the fresh feed coarsens (n decreases), which is likely coming from another crusher. In this manner, the circulating load can be related to crusher settings.

In industrial grinding process, in addition to goal of productivity maximization, other purposes of deterministic grinding circuit optimization have to satisfy the upper bound constraints on the control variables. We know that there lies a tradeoff between the throughput (TP) and the percent passing of midsize classes (MS) from the previous work of Mitra and Gopinath,2004. In deterministic optimization formulation, there are certain parameters which we will assume them as constant. But, in real life that may not be case. There are such six parameters in our industrial grinding process which are R, B, R, B are the grindability indices and grindability exponents for the rod mill (RMGI) and the ball mill (BMGI); and P, S are the sharpness indices for the primary (PCSI) and secondary cyclones (SCSI). These parameters are treated as constant in deterministic formulation. As they are going to be treated as uncertain parameters in the OUU formulation. These parameters are assumed uncertain because most of them are obtained from the regression of experimental data and thus are subject to uncertainty due to experimental and regression errors. In the next part of the section, we consider them as fuzzy numbers and solve the OUU problem by FEVM. In FEVM formulation, the uncertain parameters are considered as fuzzy numbers and the uncertain formulation is transformed into the deterministic formulation by expectation calculations for both objective function and constraints. So, the converted deterministic multi-objective optimization problem is expressed as:

Another spinning batch concentrator (Figure 10.27), it is designed principally for the recovery of free gold in grinding circuit classifier underflows where, again, a very small (<1%) mass pull to concentrate is required. The feed first flows up the sides of a cone-shaped bowl, where it stratifies according to particle density before passing over a concentrate bed fluidized from behind by back-pressure (process) water. The bed retains dense particles such as gold, and lighter gangue particles are washed over the top. Periodically the feed is stopped, the bed rinsed to remove any remaining lights and is then flushed out as the heavy product. Rinsing/flushing frequency, which is under automatic control, is determined from grade and recovery requirements.

The units come in several designs, the Semi-Batch (SB), Ultrafine (UF), and i-Con, designed for small scale and artisanal miners. The first installation was at the Blackdome Gold Mine, British Columbia, Canada, in 1986 (Nesset, 2011).

These two batch centrifugal concentrators have been widely applied in the recovery of gold, platinum, silver, mercury, and native copper; continuous versions are also operational, the Knelson Continuous Variable Discharge (CVD) and the Falcon Continuous (C) (Klein et al., 2010; Nesset, 2011).

To liberate minerals from sparsely distributed and depleting the ore bodies finer grinding than generally obtained by the conventional Rod Mill Ball Mill grinding circuits is needed. Longer grinding periods in the conventional milling processes prove too expensive mainly due to large power consumption. Stirrer mills have been tried in mineral industry with considerable success and have therefore been increasingly used. In this chapter, the theories involved in the design and operation of these mills, as established till now, are explained. Further theoretical studies and designs of the mills are still in progress for a better understanding and improved operation. Presently, the mills have been proved to be economically viable and the mineral of interest conducive to improved recovery and grade.

IMP Technologies Pty. Ltd. has recently tested a pilot-scale super fine crusher that operates on dry ore and is envisaged as a possible alternative to fine or ultra-fine grinding circuits (Kelsey and Kelly, 2014). The unit includes a rotating compression chamber and an internal gyrating mandrel (Figure 6.13). Material is fed into the compression chamber and builds until the gyratory motion of the mandrel is engaged. Axial displacement of the compression chamber and the gyratory motion of the mandrel result in fine grinding of the feed material. In one example, a feed F80 of 300m was reduced to P80 of 8m, estimated to be the equivalent to two stages of grinding. This development is the latest in a resurgence in crushing technology resulting from the competition of AG/SAG milling and the demands for increased comminution energy efficiency.

The iron oxide crystal grains in most iron ores are not evenly distributed in size. Spiral separators can therefore be used to take out the coarser iron oxide grains in the primary grinding circuit to save grinding energy and help achieve a higher iron recovery. Figure 9.14 presents a typical flow sheet for processing an oxidized ore containing about 30% Fe using a combination of spiral and SLon magnetite separators and reverse flotation. This ore is mainly composed of hematite, magnetite, and quartz, and the iron oxide crystals range in size from 0.005 to 1.0mm with an average size of about 0.05mm. The average size of the quartz crystals is approximately 0.085mm.

In the primary grinding stage of the flow sheet in Figure 9.14, the ore is first ground down to about 60% -75m and then classified into two size fractions, a coarse size fraction and a fine size fraction. The coarse size fraction is treated with spiral separators to recover part of the final iron ore concentrate. Then, drum LIMS and SLon magnetic separators are used to reject some of the coarse gangue minerals as final tailings. The magnetic products from the LIMS and SLon are sent back to the secondary ball mill for regrinding, and the milled product returns to the primary cyclone classifier.

The fine size fraction is about 90% -75m and is processed using drum LIMS separators and SLon magnetic separators in series to take out the magnetite and hematite, respectively. The magnetic products from the magnetic separators are mixed to generate the feed for reverse flotation to produce another component of the final iron ore concentrate.

The key advantage of this flow sheet lies in the fact that the spirals and SLon magnetic separators take out about 20% of the mass of the final iron concentrate and about 20% of the mass of the final tailings, respectively, from the coarse size fraction. This greatly reduces the masses being fed to the secondary ball mill and reverse flotation, thereby greatly reducing the total processing cost. From the plant results for this flow sheet, an iron concentrate containing 67.5% Fe could be produced from a run-of-mine ore containing 30.1% Fe, at a mass yield to the iron concentrate of 34.9%, an iron recovery of 78.0%, and a tailings grade of 10.2% Fe.

The first step of physical beneficiation is crushing and grinding the iron ore to its liberation size, the maximum size where individual particles of gangue are separated from the iron minerals. A flow sheet of a typical iron ore crushing and grinding circuit is shown in Figure 1.2.2 (based on Ref. [4]). This type of flow sheet is usually followed when the crude ore contains below 30% iron. The number of steps involved in crushing and grinding depends on various factors such as the hardness of the ore and the level of impurities present [5].

Jaw and gyratory crushers are used for initial size reduction to convert big rocks into small stones. This is generally followed by a cone crusher. A combination of rod mill and ball mills are then used if the ore must be ground below 325 mesh (45m). Instead of grinding the ore dry, slurry is used as feed for rod or ball mills, to avoid dusting. Oversize and undersize materials are separated using a screen; oversize material goes back for further grinding.

Typically, silica is the main gangue mineral that needs to be separated. Iron ore with high-silica content (more than about 2%) is not considered an acceptable feed for most DR processes. This is due to limitations not in the DR process itself, but the usual customer, an EAF steelmaking shop. EAFs are not designed to handle the large amounts of slag that result from using low-grade iron ores, which makes the BF a better choice in this situation. Besides silica, phosphorus, sulfur, and manganese are other impurities that are not desirable in the product and are removed from the crude ore, if economically and technically feasible.

While used sometimes on final concentrates, such as Fe concentrates, to determine the Blaine number (average particle size deduced from surface area), and on tailings for control of paste thickeners, for example, the prime application is on cyclone overflow for grinding circuit control (Kongas and Saloheimo, 2009). Control of the grinding circuit to produce the target particle size distribution for flotation (or other mineral separation process) at target throughput maximizes efficient use of the installed power.

Continuous measurement of particle size in slurries has been available since 1971, the PSM (particle size monitor) system produced then by Armco Autometrics (subsequently by Svedala and now by Thermo Gamma-Metrics) having been installed in a number of mineral processing plants (Hathaway and Guthnals, 1976).

The PSM system uses ultrasound to determine particle size. This system consists of three sections: the air eliminator, the sensor section, and the electronics section. The air eliminator draws a sample from the process stream and removes entrained air bubbles (which otherwise act as particles in the measurement). The de-aerated pulp then passes between the sensors. Measurement depends on the varying absorption of ultrasonic waves in suspensions of different particle sizes. Since solids concentration also affects the absorption, two pairs of transmitters and receivers, operating at different frequencies, are employed to measure particle size and solids concentration of the pulp, the processing of this information being performed by the electronics. The Thermo GammaMetrics PSM-400MPX (Figure 4.18) handles slurries up to 60% w/w solids and outputs five size fractions simultaneously.

Other measurement principles are now in commercial form for slurries. Direct mechanical measurement of particle size between a moving and fixed ceramic tip, and laser diffraction systems are described by Kongas and Saloheimo (2009). Two recent additions are the CYCLONEtrac systems from CiDRA Minerals Processing (Maron et al., 2014), and the OPUS ultrasonic extinction system from Sympatec (Smith et al., 2010).

CiDRAs CYCLONEtrac PST (particle size tracking) system comprises a hardened probe that penetrates into the cyclone overflow pipe to contact the stream and effectively listens to the impacts of individual particles. The output is % above (or below) a given size and has been shown to compare well with sieve sizing (Maron et al., 2014). The OPUS ultrasonic extinction system (USE) transmits ultrasonic waves through a slurry that interact with the suspended particles. The detected signal is converted into a particle size distribution, the number of frequencies used giving the number of size classes measured. Applications on ores can cover a size range from 1 to 1,000m (Smith et al., 2010).

In addition to particles size, recent developments have included sensors to detect malfunctioning cyclones. Westendorf et al. (2015) describe the use of sensors (from Portage Technologies) on cyclone overflow and underflow piping. CiDRAs CYCLONEtrac OSM (oversize monitor) is attached to the outside of the cyclone overflow pipe and detects the acoustic signal as oversize particles (rocks) hit the pipe (Cirulis and Russell, 2011). The systems are readily installed on individual cyclones thus permitting poorly operating units to be identified and changed while allowing the cyclone battery to remain in operation. Figure 4.19 shows an installation of both CiDRA systems (PST, OSM) on the overflow pipe from a cyclone.

energy use of fine grinding in mineral processing | springerlink

energy use of fine grinding in mineral processing | springerlink

Fine grinding, to P80 sizes as low as 7m, is becoming increasingly important as mines treat ores with smaller liberation sizes. This grinding is typically done using stirred mills such as the Isamill or Stirred Media Detritor. While fine grinding consumes less energy than primary grinding, it can still account for a substantial part of a mills energy budget. Overall energy use and media use are strongly related to stress intensity, as well as to media size and quality. Optimization of grinding media size and quality, as well as of other operational factors, can reduce energy use by a factor of two or more. The stirred mills used to perform fine grinding have additional process benefits, such as polishing the mineral surface, which can enhance recovery.

Fine grinding is becoming an increasingly common unit operation in mineral processing. While fine grinding can liberate ores that would otherwise be considered untreatable, it can entail high costs in terms of energy consumption and media use. These costs can be minimized by performing adequate test work and selecting appropriate operating conditions. This paper reviews fine grinding technology, research, and plant experience and seeks to shed light on ways in which operators can reduce both operating costs and the environmental footprint of their fine grinding circuit.

This paper will begin by giving an overview of fine grinding and the equipment used. It will then discuss energyproduct size relationships and modeling efforts for stirred mills in particular. The paper will go on to cover typical test work requirements, the effect of media size, and the contained energy in media. In closing, specific case studies will be reviewed.

Grinding activities in general (including coarse, intermediate, and fine grinding) account for 0.5pct of U.S. primary energy use, 3.8pct of total U.S. electricity consumption, and 40pct of total U.S. mining industry energy use. Large energy saving opportunities have been identified in grinding in particular.[1]

TableI shows a very large disparity between the theoretical minimum energy used in grinding and the actual energy used. More interestingly, a fairly large difference remains even between Best Practice grinding energy use and current energy use. This suggests that large savings in grinding energy (and associated savings in maintenance, consumables, and capital equipment needed) could be obtained by improving grinding operations.

As fine grinding is typically used on regrind applications, the feed tonnages to fine grinding circuits are small compared to head tonnages, typically 10 to 30tph. However, the specific energies are often much larger than those encountered in intermediate milling and can be as high as 60kWh/t. Total installed power in a fine grinding circuit can range from several hundred kW to several MW; for example, the largest installed Isamill has 3MW installed power.[3] This quantity is small compared to the power used by a semi-autogenous mill and a ball mill in a primary grinding circuit; a ball mill can have an installed power of up to 15MW, while installed power for a SAG mill can go up to 25MW. However, the energy used for fine grinding is still significant. Moreover, as this paper seeks to demonstrate, large energy reduction opportunities are frequently found in fine grinding.

Grinding can be classified into coarse, intermediate, and fine grinding processes. These differ in the equipment used, the product sizes attained, and the comminution mechanisms used. The boundaries between these size classes must always be drawn somewhat arbitrarily; for this paper, the boundaries are as given in TableII. As shown in the table, coarse grinding typically corresponds to using an AG or SAG mill, intermediate grinding to a ball mill or tower mill, and fine grinding to a stirred mill such as an Isamill or Stirred Media Detritor (SMD). Of course, various exceptions to these typical values can be found.

In fine grinding, a material with an F80 of less than 100m is comminuted to a P80 of 7 to 30m. (P80s of 2m are at least claimed by equipment manufacturers.) The feed is typically a flotation concentrate, which is reground to liberate fine particles of the value mineral.

The three modes of particle breakage are impact; abrasion, in which two particles shear against each other; and attrition, in which a small particle is sheared between two larger particles or media moving at different velocities. In fine grinding, breakage is dominated by attrition alone.[4] In stirred mills, this is accomplished by creating a gradient in the angular velocity of the grinding media along the mills radius.

Fine grinding is usually performed in high-intensity stirred mills; several manufacturers of these stirred mills exist. Two frequently used stirred mills include the Isamill, produced by Xstrata Technology, and the SMD, produced by Metso (Figure1). A third mill, the KnelsonDeswik mill (now the FLS stirred mill), is a relative newcomer to the stirred milling scene, having been developed through the 1990s and the early 2000s.[5] In all these mills, a bed of ceramic or sand is stirred at high speed. Ceramic media sizes in use range from 1 to 6.5mm.

The Isamill and the SMD have very similar grinding performance. Grinding the same feed using the same media, Nesset et al.[7] found that the Isamill and SMD had very similar specific energy use. Gao et al.[8] observed that an Isamill and SMD, grinding the same feed with the same media, produced very similar product particle size distributions (PSDs). This similarity in performance has also been observed in other operations.

Nevertheless, there are important differences. In the Isamill, the shaft is horizontal and the media are stirred by disks, while in the SMD, the stirring is performed by pins mounted on a vertical shaft. In an SMD, the product is separated from the media by a screen; the Isamill uses an internal centrifugation system. This means that the screens in an SMD constitute a wear part that must be replaced, while for the Isamill, the seals between the shaft and body constitute important wear parts. Liner changes and other maintenance are claimed by Xstrata Technology to be much easier than in an SMD: While an SMDs liner is removed in eight parts, the Isamills liner can be removed in two pieces, with the shell sliding off easily.[3] The KnelsonDeswik mill is top stirred and can therefore be considered to be similar to an SMD.[5]

An important difference among the Isamill, the SMD, and the KnelsonDeswik mill is that of scale. The largest Isamill installed at time of writing had 3MW of installed power; an 8MW Isamill is available, but appears not to have yet been installed.[3] The largest SMD available has 1.1MW of installed power; one 1.1-MW SMD has been installed. The next largest size SMD has 355kW of installed power.[6] Thus, several SMDs are often installed for a fine grinding circuit, while the same duty would be performed by a single Isamill. SMDs are typically arranged in series, with the product of one becoming the feed for the other. This has the advantage that each SMD in the line can have its media and operating conditions optimized to the particle size of its particular feed. The largest installed power in a KnelsonDeswik mill is 699kW[5]; this places it in an intermediate position between the 355-kW and 1.1-MW SMDs.

In 2012, FLSmidth reported that it had acquired the KnelsonDeswik mill; the mill is now known as the FLSmidth stirred mill. An FLSmidth stirred mill will be installed to perform a copper concentrate regrind in Mongolia.[9] It is speculated that the mill will continue to be scaled up under its new owners to allow it to effectively compete against the SMD and Isamill.

Gravity-induced stirred (GIS) mills include the Tower mill, produced by Nippon Eirich, and the Vertimill, produced by Metso. Grinding to below 40m in GIS mills or ball mills is usually not recommended. In their product literature, Metso give 40m as the lower end of the optimal P80 range for Vertimills.[6] At lower product sizes, both tower mills and ball mills will overgrind fines. At Mt. Isa Mines, a GIS mill fed with material of F80 approximately 50m lowered the P80 size by only 5 to 10m, at the same time producing a large amount of fines.[10] Similarly, in ball mills, it is known that grinding finer than approximately 40m will result in overgrinding of fines as well as high media consumption. However, it must be noted that the product size to which a mill can efficiently grind depends on the feed material, the F80, and media type and size. A Vertimill has been used to grind to sizes below 10m.[11]

The phenomenon of overgrinding is largely the result of using media that are too large for the product size generated. The smallest ball size typically charged into ball mills and tower mills is inch (12.5mm), although media diameters as small as 6mm have been used industrially in Vertimills.[11]

In a laboratory study by Nesset et al.,[7] a GIS mill charged with 5-mm steel shot, and with other operating conditions similarly optimized, achieved high energy efficiencies when grinding to less than 20m. This appears to qualitatively confirm the notion that fine grinding requires smaller media sizes. In the case of the Nesset study, the power intensity applied to the laboratory tower mill was lowthat is, the shaft was rotated slowly in order to obtain this high efficiency, leading to low throughput. This suggests that charging GIS mills with small media may not be practicable in plant operation.

Millpebs have been used as grinding media to achieve fine grinding in ball mills. These are 5- to 12-mm spherical or oblong cast steel pellets, charged into ball mills as a replacement of, or in addition to, balls. While Millpebs can give significantly lower energy use when grinding to finer sizes, they also can lead to high fines production and high media use.

Millpebs were tested for fine grinding at the Brunswick concentrator. The regrind ball mills at the concentrator used 25-mm slugs to produce a P80 of 28m. In one of the regrind mills, the slugs were replaced by Millpebs; these were able to consistently maintain a P80 of 22m while decreasing the power draw by 20pct. However, media use increased by 50pct and the production of fines of less than 16m diameter increased by a factor of 5.[12] The observed drop in specific energy may be due to the fact that Millpebs had smaller average diameters than the slugs and so were more efficient at grinding to the relatively small product sizes required. It is therefore unclear whether the performance of Millpebs would be better than that of conventional 12-mm steel balls. To the best of the authors knowledge, no performance comparison between Millpebs and similarly sized balls has been performed.

A host of other technologies exist to produce fine grinding, including jet mills, vibrating mills, roller mills, etc. However, none of these technologies has reached the same unit installed power as stirred mills. For example, one of the largest vibrating mills has an installed power of 160kW.[13] Therefore, these mills are considered as filling niche roles and are not treated further in this review. A fuller discussion of other fine grinding technologies can be found in a review by Orumwense and Forssberg.[14]

Neese et al.[15] subjected 50- to 150-m sand contaminated with oil to cleaning in a stirred mill in the laboratory. The mill operated at low stress intensities: A low speed and small-size media (200- to 400-m quartz or steel beads) were used. These conditions allowed the particles to be attrited without being broken. As a result, a large part of the oil contaminants was moved to the 5-m portion of the product. This treatment may hold promise as an alternative means of processing bituminous sands, for example, in northern Alberta.

The Albion process uses ultrafine grinding to enhance the oxidation of sulfide concentrates in treating refractory gold ores.[16] In the process, the flotation concentrate is ground to a P80 of 10 to 12m. The product slurry is reacted with oxygen in a leach tank at atmospheric pressure; limestone is added to maintain the pH at 5 to 5.5. The leach reaction is autothermal and is maintained near the slurry boiling point. Without the fine grinding step, an autoclave would be required for the oxygen leaching process. It is hypothesized that the fine grinding enhances leach kinetics by increasing the surface area of the particles, as well as by deforming the crystal lattices of the particles.

Numerous researchers, for example, Buys et al.,[17] report that stirred milling increases downstream flotation recoveries by cleaning the surface of the particles. The grinding media used in stirred mills are inert, and therefore corrosion reactions, which occur with steel media in ball mills, are not encountered. Corrosion reactions change the surface chemistry of particles, especially with sulfide feeds, and hamper downstream flotation.

Further increases in flotation recoveries are obtained by limiting the amount of ultrafine particles formed; stirred mills can selectively grind the larger particles in the feed with little increase in ultrafines production. Ultrafine particles are difficult to recover in flotation.

In intermediate grinding to approximately 75m, the Bond equation (Eq. [1]) is used to relate feed size, product size, and mechanical energy applied. Below 75m, correction factors can be applied to extend its range of validity.[4]

No general work index formula governing energy use over a range of conditions, like the Bond equation for intermediate grinding, has yet been found for the fine grinding regime. Instead, the work-to-P80 curve is determined in the laboratory for each case. The energy use usually fits an equation of the form

Signature plot (specific energy vs P80 curve) for Brunswick concentrator Zn circuit ball mill cyclone underflow; F80=63m. The plots give results for grinding the same feed using different mills and media. After Nesset et al.[7]

Values for the exponent k have been found in the range 0.7 to 3.5, meaning that the work to grind increases more rapidly as grind size decreases than in intermediate grinding. The specific energy vs product size curve has a much steeper slope in this region than in intermediate grinding.

The values of k and A are specific to the grinding conditions used in the laboratory tests. Changes in feed size, media size distribution, and in other properties such as media sphericity and hardness can change both k and A, often by very large amounts. Media size and F80 appear to be the most important determinants of the signature plot equation.

The connections (if any) between k and A and various operating conditions remain unknown. Because of the relatively recent advent of stirred milling in mineral processing, fine grinding has not been studied to the same extent as grinding in ball mills (which of course entail much larger capital and energy expenditures in any case). One of the research priorities in the field of stirred milling should be the investigation of the effects of F80 and media size on the position of the signature plots. If analogous formulas to the Bond ball mill work formula and the Bond top ball size formula can be found, the amount of test work required for stirred milling would be greatly reduced.

Larson et al.[19] found that when specific energy is plotted against the square of the percent particles in the product passing a given size (a proxy for particle surface area), a straight line is obtained. This is demonstrated in Figure3.

In contrast to the conventional signature plot, this function gives zero energy at the mill feed. It is therefore hypothesized that if a squared function plot is obtained by test work for one feed particle size, the plot for another feed particle size can be obtained simply by changing the intercept of the line while keeping the slope the same. Therefore, the squared function plot allows the effect of changes in both F80 and P80 to be modeled.

While the Squared Function Plot is intriguing, experimental validation of its applicability has not yet been published. It nevertheless remains an interesting topic for further investigation and if validated may be used in the future as an alternative measure of specific energy.

A similar analysis has been performed by Musa and Morrison,[21] who developed a model to determine the surface area within each size fraction of mill product. They defined a marker size below which 70 to 80pct of the product surface area was contained; the marker size thus served as a proxy for surface area production. Specific energy use was then defined as kWh of power per the tonne of new material generated below the marker size. Musa and Morrison found that by defining specific energy in this way, it was possible to accurately predict the performance of full-scale Vertimills and Isamills from laboratory tests.

Blecher and coworkers[22,23] found that stress intensity combines the most important variables determining milling performance. Stress intensity for a horizontal stirred mill, with media much harder than the mineral to be ground, is defined as in Eq. [4].

Note that the stress intensity is strongly sensitive to changes in media diameter (to the third power), is less sensitive to stirrer tip speed (to the second power), and is relatively insensitive to media and slurry density.

For vertical stirred mills such as the SMD and tower mill, both SIs and SIg are non-zero. For horizontal stirred mills such as the Isamill, net gravitational SI is zero due to symmetry along the horizontal axis. Therefore, for horizontal stirred mills, only SIs need be taken into consideration.

Kwade and coworkers noted that, at a given specific energy input, the product P80 obtainable varies with stress intensity and passes through a minimum. Product size at a given energy input can be viewed as a measure of milling efficiency; therefore, milling efficiency reaches a maximum at a single given stress intensity. This idea was experimentally validated by Jankovic and Valery (Figure 4).[25]

The stress intensity is defined by parameters that are independent of mill size or type. According to Jankovic and Valery,[25] once the optimum SI has been determined in one mill for a given feed, the same SI should also be the point of optimum efficiency in any other mill treating that feed. Therefore, the optimum SI need only be determined in one mill (e.g., a small test mill); the operating parameters of a full-scale mill need only be adjusted to produce the optimum SI.

Stress frequency multiplied by stress intensity is equal to mill power; therefore, stress intensity could in theory be used to predict mill specific energy. However, to the authors knowledge, a comprehensive model linking stress intensity, stress frequency, and specific energy has not yet been developed. Therefore, there is not yet any direct link between stress intensity and specific energy.

The definition of SIs as given in Eq. [4] is valid only for cases where the grinding media are much harder than that of the material ground (for example, the grinding of limestone with glass beads). Becker and Schwedes[26] determined that, in a collision between media and a mineral particle, the fraction of energy transferred to the product is given by Eq. [6]:

To maintain high efficiency in milling, the media must be chosen so as to be much harder (higher Youngs modulus) than the product material, keeping E p,rel close to unity. Where the Youngs modulus of the product is similar to that of the media, much of the applied energy goes into deformation of the media instead of that of the particle to be ground. The energy used to deform the media is lost, lowering the amount of energy transferred to the product. This fact explains why steel media, with a relatively low Youngs modulus, tend to perform poorly in stirred milling, even though the media are much more dense than silica or alumina media.

The previous sections indicated that stress intensity is independent from individual millsi.e., the optimal stress intensity when using Mill A will also be the optimal stress intensity when using Mill B. However, this does not seem to be the case when actually scaling up mills.

Four-liter Isamills are commonly used for grindability test work. It can be assumed that operating parameters of the test mill (including media type, media size, and slurry density) are adjusted so far as possible to give the optimum SI. These parameters are then used in the full-scale mill as well. However, the 4-L test mills have a tip speed of approximately 8m/s, while full-scale Isamills have tip speeds close to 20m/s.[27] If the same media size, media density, and slurry density are used in the test mill as in the full-scale mill, the stress intensity of the full-scale mill will be approximately 6.25 times larger than that of the test mill. This implies that the full-scale mill is operating outside of the optimum SI and will be grinding less efficiently. That is to say that the operating point of the full-scale mill will be above the signature plot determined by test work.

In reality, however, the operating points of full-scale stirred mills are generally found to lie on the signature plots generated in test work.[19] Therefore, the full-scale mills and test mills have the same milling efficiency, even though the full-scale mill operates at a different stress intensity than the test mill.

This question remains unresolved. One possible answer arises from the observation that two of the P80 vs SI curves in Figure4 appear to have broad troughs, covering almost an order of magnitude change in SI. In this case, even a sixfold increase in SI might not create a noticeable difference in performance, considering experimental and measurement error.

Product size vs stress intensity at three different specific energies for a zinc regrind. Note optimum stress intensity at which the lowest product size is reached. Figure used with permission from Jankovic and Valery[25]

The SMD test unit appears from photographs to have a bed depth of around 30cm, while the full-scale SMD355 has a bed depth of approximately one meter. This represents a change in the gravitational stress intensity of almost two orders of magnitude. As has been previously noted, however, laboratory and full-scale SMDs scale-up with a scale-up factor of approximately unity, with no apparent change in the optimum stress intensity. This observation suggests that the gravitational stress intensity, SIg, is unimportant in SMDs compared to the stirring stress intensity, SIs. By contrast, in GIS mills, where full-size units have bed depths of ten meters or more, gravitational stress intensity can be expected to be much more important in full-size units than in test units, adding a complicating factor to GIS mill scale-up.

Factorial design experiments were performed by Gao et al.[28] and Tuzun and Loveday[29] to determine the effect of various operating parameters on the power use of laboratory mills. Power models were determined giving the impact of different parameters as power equations with linear and nonlinear terms. The derived models did not appear to be applicable to mills other than the particular laboratory units being studied.

In ball milling, the Bond ball mill work index can be used to determine specific energy at a range of feed and product sizes. The Bond top size ball formula can be used to estimate the media size required. No such standard formulas exist in fine grinding. Energy and media parameters must instead be determined in the laboratory for every new combination of operating conditions such as feed size, media size, and media type.

For the Isamill, test work is usually performed with a 4-L bench-scale Isamill. Approximately 15kg of the material to be ground is slurried to 20pct solid density by volume. The slurry is then fed through the mill and mill power is measured. The products PSD is measured, additional water is added if needed, and the material is sent through the mill again. This continues until the target P80 is reached; typically, there will be 5 to 10 passes through the mill. The test work will produce a signature plot and media consumption data as the deliverables.

In contrast to laboratory-scale testing for ball mills and AG/SAG mills, test work results for stirred mills can be used for sizing full-size equipment with a scale-up factor close to one. Larson et al.[19,20] found a scale-up factor for the Isamill of exactly 1, while Gao et al.[8] imply that the scale-up factor for SMDs is 1.25.

A common error in test work is using monosize media (e.g., fresh 2-mm media loaded into in the mill) as opposed to aged media with a distribution of particle sizes. The aged media will grind the smaller feed particles more efficiently. Therefore, using fresh media will give a higher specific energy than in reality.[30]

Another pitfall is coarse holdup in the mill. If the mill is not sufficiently flushed, coarse particles will be kept inside the mill. The mill product then appears finer than it in reality is. This leads to lower estimates of specific energy than reality.[19]

In ball milling, the product particle size distribution (PSD) can usually be modeled as being parallel to the feed PSD on a log-linear plot.[4] When grinding to finer sizes in ball mills, the parallel PSDs mean that large amounts of ultrafine particles are produced. This consumes a large amount of grinding energy while producing particles which are difficult to recover in subsequent processing steps such as flotation.

As shown in the figure, at the left end of the graph, the product PSD is very close to the feed PSD; at the right, the two PSDs are widely spaced. This indicates that the mill is efficiently using its energy to break the top size particles and is spending very little energy on further grinding of fine particles. Thus, the overall energy efficiency of the fine grinding can be expected to be good. As a bonus, the tighter PSD makes control of downstream processes such as flotation easier.

In an experimental study, Jankovic and Sinclair subjected calcite and silica to fine grinding in a laboratory pin stirred mill, a Sala agitated mill (SAM), and a pilot tower mill. The authors found that for each mill, the PSD of the product was narrower (steeper) than that of the feed. In addition, when grinding to P80s below approximately 20m in any of the three mills tested, the PSD became more narrow (as measured by P80/P20 ratio) as the P80 decreased. (When the width of the PSD was calculated using an alternative formula, the PSD was only observed to narrow with decreasing P80 when using the pin stirred mill.) The authors concluded that the width of the PSD was strongly affected by the material properties of the feed, while not being significantly affected by the media size used.[32]

In stirred milling, the most commonly used media are ceramic balls of 1 to 5mm diameter. The ceramic is usually composed of alumina, an alumina/zirconia blend, or zirconium silicate. Ceramic media exist over a wide range of quality and cost, with the lower quality/cost ceramic having a higher wear rate than higher quality/cost ceramic. Other operations have used sand as media, but at the time of writing, only two operations continue to use sand.[8,27,33] Mt Isa Mines has used lead smelter slag as media; however, it is now using sand media.[10,27] Mt Isa is an exception in its use of slag, as a vast majority of operations do not have a smelter on-site to provide a limitless supply of free grinding media. However, in locations where slag is available, it should be considered as another source of media.

Media use in fine grinding is considered to be proportional to the mechanical energy applied. Typical wear rates and costs are given in TableIII and Figure6; these figures can of course vary significantly from operation to operation.

Contained energy refers to the energy required to produce and transport the media, and is distinct from the mechanical (electrical) energy used to drive the mill. Hammond and Jones estimated the contained energy in household ceramics (not taking account of transportation).[39] Hammond and Jones estimates range from 2.5 to 29.1MJ/kg, with 10MJ/kg for general ceramics and 29MJ/kg for sanitary ceramics. Given that ceramic grinding media require very good hardness and strength, especially compared to household ceramics, it is appropriate to estimate its contained energy at the top end of Hammond and Jones range, at 29MJ/kg.

Using 29MJ/kg for the contained energy of ceramic media and a wear rate of 35g/kWh of mechanical energy gives a contained energy consumption of 0.28kWh contained per kWh of mechanical energy applied. A wear rate of 7g/kWh gives a contained energy consumption of 0.06kWh contained per kWh of mechanical energy applied. Therefore, 6 to 20pct of the energy use in fine grinding using ceramic media can be represented by contained energy in the grinding media itself.

Sand media have much lower contained energy than ceramic media as the media must simply be mined or quarried rather than manufactured. Hammond and Jones report a contained energy of 0.1MJ/kg. Blake et al.[36] reported that switching a stirred mills media from sand to ceramic results in a mechanical energy savings of 20pct. Therefore, using sand rather than ceramic media would produce savings in contained energy, but would cost more in mechanical energy. Likewise, Davey[40] suggests that poor-quality media will increase mechanical energy use in stirred milling. It is speculated that this is due to the lower sphericity of sand media. On the other hand, the work of Nesset et al.[7] suggests that the energy use between ceramic and sand media of the same size is the same. Slag media, where a smelter is on-site, would probably have the lowest contained energy consumption of the different media types. There is very little transportation, and for accounting purposes, almost no energy has gone into creating the media as the granulated slag is a by-product of smelter operation.

Becker and Schwedes[41] point out that with poor-quality media, a significant part of the product will consist of broken pieces of media, which will affect the measured product PSD. Clearly, more information on the relationships between contained energy in media and media wear rates is desirable.

Of the different operating parameters for stirred mills, media size probably has the biggest influence on overall energy consumption. The appropriate media size for a mill appears to be a function of the F80 and P80 required. The grinding media must be large enough to break up the largest particles fed to the mill and small enough to grind the material to the product fineness desired. As demonstrated by the experience of Century mine, an inappropriate media size choice can result in energy consumption double that of optimum operation.[8]

In their laboratory study, Nesset et al.[7] varied a number of operating parameters for stirred mills and identified media size as having the largest impact on energy use. It was also noted that the trials which produced the sharpest product PSD were also the ones which resulted in the lowest specific energy use.

Gao et al.[8] report that at Century mine, the grinding media in SMDs performing regrind duty were changed from 1 to 3mm. This resulted in a drop in energy use of approximately 50pct; the signature plot shifted significantly downward (Figure7).

Figure8 shows the product PSD for laboratory SMD tests using 1- and 3-mm media. The PSD for the test using 1-mm media shows that the SMD produced a significant amount of fines (20pct below 4m). The mill also had difficulty breaking the top size particlesthe 100pct passing size appears to be almost the same for both the feed and the product. In contrast, the PSD using 3-mm media shows less fines production (20pct below 9m) and effective top size breakage, with all the particles above 90m broken. This is in line with the observation of Nesset et al.[7] that low energy use is associated with tight product size distributions.

Gao et al.[38] tested copper reverberatory furnace slag (CRFS, SG 3.8) and heavy media plant rejects (HMPR, SG 2.4) in a laboratory stirred mill at two sizes: 0.8/+0.3mm, and 1.7/+0.4mm. For both CRFS and HMPR, the smaller size media gave a lower specific energy than the larger size media. At the same size, both CRFS and HMPR had similar specific energy use. However, the CRFS ground the material much faster than HMPR. Possibly, this was due to its higher density.

Data on F80, P80, and media size were compiled from the literature in order to allow benchmarking against existing operations. The sources are listed in Table IV. F80 and P80 were plotted against media size; the results are given in Figure9.

F80 plotted against media size (blue diamonds); P80 plotted against media size (red crosses). Century UFG=Century ultrafine grind; Century Regr.=Century regrind. Data are taken from Case studies table (Color figure online)

It can be seen from the figure that as the P80 achieved decreases, the media size does as well, from 3mm to achieve 45m to 1mm to achieve under 10m. The F80 decreases with media size in a similar way, from 90m at 3mm to 45m at 1mm. Dotted lines have been added to Figure7 to define the region of operation of mills; these delimit a zone in which the stirred mill can be expected to operate efficiently.

In general, for a particular media size, limits on both F80 and P80 must be respected. For example, the figure suggests that a mill operating with an F80 of 100m should use 3-mm media, while a mill grinding to below 10m would need to use 1-mm media. To reduce a feed of 90m F80 to 10m P80, Figure9 suggests that comminution be done in two stages (two Isamills or SMDs in series) for optimal efficiency. The first stage would grind the feed from 90m to perhaps 45m using 3-mm media, while the second would grind from 45 to 10m using 1- or 2-mm media.

A number of opportunities exist to reduce the energy footprint of fine grinding mills. There are no general formulas, such as the Bond work formula and Bond top size ball formula in ball milling, to describe the performance of stirred mills. Therefore, improvement opportunities must be quantified by performing appropriate test work.

In addition to obtaining the signature plot, the specific energy as a function of new surface area should be determined during test work. This could be done either by the method of Larsen or by that of Musa and Morrison. Defining specific energy as a function of new surface area may constitute a superior means of predicting the performance of full-scale mills, as opposed to defining specific energy as a function of feed tonnage.

Media size should be chosen with care. It is recommended that test work be done with several media sizes in order to locate the stress intensity optimum. Media size can be benchmarked against other operations using Figure9.

There are indications that lower-quality media, apart from degrading faster, require more mechanical energy for grinding due to factors such as lower sphericity. It is recommended to perform test work using media of different quality to determine the effect of media quality on energy use. Slag and sand media may also be considered. Subsequently, a trade-off study involving media cost, electricity cost, improvement in energy efficiency, and contained energy in media should be performed to identify the best media from an economic and energy footprint standpoint.

D. Rahal, D. Erasmus, and K. Major: KnelsonDeswick Milling Technology: Bridging the Gap Between Low and High Speed Stirred Mills, Paper presented at the 43rd Canadian Mineral Processors Meeting, Ottawa, 2011.

Metso: Stirred milling: Vertimill grinding mills and Stirred Media Detritor (product brochure), 2013, available at http://www.metso.com/miningandconstruction/MaTobox7.nsf/DocsByID/F58680427E2A748F852576C4005210AC/$File/Stirred_Mills_Brochure-2011_LR.pdf, accessed April 21, 2013.

J. Nesset, P. Radziszewski, C. Hardie, and D. Leroux: Assessing the Performance and Efficiency of Fine Grinding Technologies, Paper presented at the 38th Canadian Mineral Processors Meeting, Ottawa, 2006.

FLSmidth: Acquisition enhances our precious metals offerings, 2012, FLSmidth eHighlights April 2012, available at http://www.flsmidth.com/en-US/eHighlights/Archive/Minerals/2012/April/Acquisition+enhances+our+precious+metals+offerings, accessed 17 April 2013.

S. Buys, C. Rule, and D. Curry: The Application of Large Scale Stirred Milling to the Retreatment of Merensky Platinum Tailings, Paper presented at the 37th Canadian Mineral Processors Meeting, Ottawa, 2005.

D. Curry, M. Cooper, J. Rubenstein, T. Shouldice, and M. Young: The Right Tools in the Right Place: How Xstrata Nickel Australasia Increased Ni Throughput at Its Cosmos Plant, Paper presented at the 42nd Canadian Mineral Processors conference, Ottawa, 2010.

G. Davey: Fine Grinding Applications Using the Metso Vertimill Grinding Mill and the Metso Stirred Media Detritor (SMD) in Gold Processing, Paper presented at the 38th Canadian Mineral Processors Meeting, Ottawa, 2006.

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best energy consumption

best energy consumption

When it comes to achieving the best energy consumption, what are the key factors a cement producer needs to address? In this article, extracted from the newly published Cement Plant Environmental Handbook (Second Edition), Lawrie Evans presents a masterclass in understanding and optimising cement plant energy consumption. By Lawrie Evans, EmCem Ltd, UK.

As control of sources, generation, distribution and consumption of energy is central to many current world issues, controlling the industrys energy footprint is a matter of intense interest to governments. This is recognised in such initiatives as ISO 50001, the World Business Council for Sustainable Developments Cement Sustainability Initiative, Energy Star in the USA, PAT in India and CO2 taxes/trading in Europe and in other countries.

For the cement industry, there are three main drivers to energy consumption: electrical power fuel customer demand for high-strength products that require a significant proportion of high-energy clinker as a component.

For the producer, these factors have a significant influence on cost competitiveness, usually accounting for over 50 per cent of total production costs, so that accurately and continuously monitoring energy usage must be a way of life for any producers technical team. The introduction of CO2 taxes in Europe and elsewhere adds a further twist to the story. For major groups, especially, decisions made in balancing maintenance, investments, operations and purchasing requirements all have to take into account the impact on their energy footprint.

Globally a cement major such as Italcementi consumes annually some 6000GWh of power and 35,500,000Gcal of heat for a total of 5Mtpe. This is the same total energy as consumed by approximately 1.6m Italians or 0.6m Americans per year. For fuel-related energy costs, the worldwide industry has largely moved to efficient preheater/precalciner processes and has found many options to switch to cheaper fuels, with the global drive to alternative fuels still proceeding. For electrical energy, options to reduce unitary costs are much more limited in scope. Most countries still have power generation/distribution systems that are effective monopolies and the cement producers cost control capability is usually limited to selecting the appropriate contract and taking opportunities offered in lower-cost off-peak power tariffs, where they exist.

Figure 1 illustrates the wide variation in the cost of power across 14 countries. The average country cost of electrical power at an industrial level varies enormously. When the added complexity of on and off peak power costs, interruption clauses, supply charges versus energy charges, etc, are added, the evaluation of the benefits of energy saving investment can become very complex. Typical cement plant power costs can range from EUR39 to EUR170/MWh.

The most important first step in controlling energy consumption is to be aware of the relative importance of the process areas where most energy is consumed. Figure 2 shows a typical breakdown of electrical energy consumption at a cement plant. The most obvious area for attention is that of grinding, both raw and cement. In either case, grinding is, by design, a very inefficient process.

The ball mill has been the industrys workhorse for over a century and despite its estimated meagre four per cent efficiency, little has changed over the years other than increases in the wear resistance of mill internals and the scale of the equipment. The addition of closed circuiting and progressively higher efficiency separators has improved cement product quality and produced higher outputs for a given mill size, but the case for adding or upgrading separators on energy saving alone has proved to be poor, unless the products are >4000Blaine. Starting from the 1970s, a new generation of mills appeared. Vertical mills (see Figure 3) were common for solid fuel grinding, generally with spring-loaded rollers. The principle of the new generation of vertical mill was to direct higher pressure from the grinding element to the material bed using hydraulic systems. From this approach the roller press, CKP (pre-grind vertical rollers) and Horomill all developed.

The gas-swept vertical mill quickly became the raw mill of choice. Grinding energy was approximately 50 per cent of the ball mill and the drying capabilities allowed direct processing of materials of up to 20 per cent moisture content. The main energy issue was the high power consumption of mill fans, with pressure drops of 100mbar not uncommon with high nozzle ring velocities (>70m/s) and internal mill circulating loads of >1000 per cent. Manufacturers have countered this generally satisfactorily with pressure drops reduced by lower nozzle ring velocities and the addition of external spillage elevator recirculation systems plus higher-efficiency separators.

Better seal designs for mill roller assemblies and pull rods have reduced the inevitable inleaking air issue and its impact on power consumption. However, it remains a design where issues of wear and reliability are more challenging than for ball mills, and these issues have not diminished with increased scale. For raw grinding with relatively dry raw materials, the combination of the roller press and V separator is a viable alternative with far lower mill fan power.

For cement grinding, the technology development away from ball mills has taken a different route. The development of roller presses in the 1980s took advantage of the benefits of higher-pressure grinding and many presses were retrofitted to ball mills as pregrinders. The main benefit was seen at lower Blaines as the first generation of presses suffered from stability problems when attempts were made to grind more finely by recirculating separator rejects. These problems are now largely resolved and the combination of a V and third-generation dynamic classifier separators together with a roller press can produce finished cement with high energy efficiency.

The Horomill and CKP systems have also enjoyed some market success and have provided good energy efficiency levels compared to ball mills. The vertical mill option has been slower to enter the cement grinding market. Grinding bed stability problems offered a challenge which the major manufacturers battled with, until finally a significant number of mills began to be installed in the late 1990s, and this has multiplied in the past decade. However, in pure energy efficiency terms, the benefit of grinding power reduction is countered by the very high power required by mill fans. In addition, the absence of the heat generated in a ball mill and the high volume of air required by the vertical mill have required the provision of waste heat from cooler exhausts and/or auxiliary furnaces to dry raw materials and achieve a limited dehydration of gypsum.

A typical comparison of three competing technologies is given in Table 1, demonstrating that an efficient ball mill/third-generation separator, CKP/ball mill/third-generation separator and vertical mill on a typical 4000Blaine limestone cement show little overall difference in energy consumption. Considering the higher capital cost, and more demanding maintenance and operating regime, there is no clear energy case to favour some of the modern variants.

Even for solid fuel grinding, there has been a minor trend back to ball mills. This is most evident for petcoke grinding, where the demand for very low residues, and the very hard and sometimes abrasive nature of high-sulphur cokes has resulted in ball mill selection.

Many of the grinding design issues, which are still under debate, are usually very clear in other areas of process selection: high-efficiency process fans and low-pressure drop preheaters adequately-sized bag filters for the main exhaust to avoid high pressure drops and poor bag life avoidance of pneumatic transport systems low-energy raw meal homogenisation silos.

The main continued discussions are those of two- or three-fan systems for the raw mill/kiln or single filter for kiln and cooler, precipitator or bag filter for the cooler exhaust and two or three tyre kiln. For a bag filter on a separate cooler the main equipment energy efficiency issue is the air-to-air heat exchanger, but this is often substituted with a water spray in the cooler or more recently, by using a ceramic filter capable of operating at above 400C.

Finally, in design terms, the most difficult decision is to avoid overdesign by applying too many safety factors. Post-commissioning audits often uncover a high contribution to poor energy efficiency from under-run equipment operating where it cannot perform efficiently.

In normal operations maintenance also plays a major part in ensuring energy efficiency. The impact of poor plant reliability upon overall electrical energy consumption is often under-estimated. In the kiln area, 100 short/medium stops (30 minutes to eight hours) per year can cost up to 5kWh/t clinker. The avoidance of inleaking air, correct alignment of motors, stopping compressed air leaks, etc are all part of the value of good maintenance.

In the key area of grinding there are important factors to control. For ball mills, ball charge level, lining and diaphragm condition must be monitored and maintained in near-optimum condition. Mill stops, defined as mill motor off, and measured by mean time between failures (mtbf), are frequently poorly recorded and the resolution of underlying issues is frequently not addressed.

Instability, where ball mill feed is stopped and the mill ground out, is also infrequently recorded or acted upon. When it comes to mill control, operators rarely concentrate on pushing mill production when the kiln is regarded as the key. Expert systems on mills should be universal and well tuned.

Grinding aids can give benefits of 5-15 per cent in production but need to be continuously evaluated for cost effectiveness. Unfortunately, their cost has risen more rapidly than the cost of energy in recent years and the economic balance has to be re-evaluated. The benefit of aids on cement flowability has to be considered, along with the added scope for reduction of cement clinker content with some modern additives. Correct timing on the maintenance of a first chamber cement mill lining and the successful implementation of an expert system on a cement mill both offer benefits in terms of power consumption (see case studies panel). Accurate process measurements are also key to energy saving opportunities. Air compressors are another area for attention. Often, these are multiple units operating on a cycle of on- and off-load. Replacement of one (of three) with a variable-speed type (see Table 2) can provide rapid payback. Even lighting and buildings offer excellent opportunities for power savings. Table 3 shows the 40-80 per cent energy savings that can be achieved by simply replacing old lighting systems. Buildings such as the new Italcementi Group Research and Innovation Centre (i.lab) in Bergamo, Italy, demonstrate that good building design creates significant savings.

A major change has occurred in the last 20 years in the area of in-house generation of electrical energy by cement manufacturers, most significantly using waste heat recovery (WHR) from the pyroprocessing line. Figure 4 shows the areas suited to heat recovery for power generation, and WHR technology is already applied to preheater and cooler exhausts.

The modern technology originated in Japan in the 1980s, where high power prices and large-scale operations combined to produce useful economic returns, with most applications using steam boilers at the preheater exhaust. Little further development happened outside Japan until the turn of the century, when a combination of lower capital cost, Chinese equipment, and the idea to improve recovery by splitting cooler exhausts into higher and lower temperature streams combined to offer the paybacks necessary for the technology to expand, first inside China and then beyond.

The results of WHR have been impressive, eg, with the 19MW net achieved from a combined installation on two five-stage precalciner kilns (5500tpd and 7500tpd) in Thailand being typical. Options for the technology are evolving with other thermodynamic cycles being applied: steam Rankine cycle with various enhancements the most widely applied technology organic Rankine cycle a variety of organic fluids applied and favoured at lower gas temperatures Kalina (ammonia/water) cycle supercritical CO2 cycle.

There are also further developments which can increase the power recovered, including recycling the lower temperature cooler exhaust, meal bypassing preheater stages to boost exit temperatures and the use of alternative fuels and excess air, also to boost preheater exit temperature and energy recovery. Other options for power generation can use the land owned by the cement plant for raw material reserves. These include wind farms photovoltaics, concentrated solar panels or growing and burning biomass either to boost power in a WHR system or for use in an internal, stand-alone power generation plant.

pinion shaft - an overview | sciencedirect topics

pinion shaft - an overview | sciencedirect topics

The typical IGC pinion shaft arrangement, having the impellers outboard of the seals and bearings, allows access for the application of VIGV mechanisms at the inlet of each stage (Fig. 4.5A), if desired, and similar accessibility exists for VDVs (Fig. 4.5B) with similar actuation mechanisms. In both cases, a single actuator rotates a control ring, which subsequently rotates each IGV or VDV in unison via identical cam features. Variable geometry can be used to increase the peak performance or operating range of a compressor, as is demonstrated in a later section. Because of the relative ease of access, VIGVs and VDVs are significantly more cost effective to implement on IGCs than with typical inline compressors. However, while this accessibility exists to varying degrees for every stage in an IGC, it is common to only implement VIGVs and/or VDVs for the first stage.

The output shaft is connected to the input pinion shaft of the gearbox by a high speed coupling shaft. This has flexible diaphragm elements and so allows some lateral misalignment between the GT and gearbox but, more importantly, allows axial tension (prestretch) to be incorporated when the unit is in the cold (ambient temperature) stationary condition. This axial prestretch is necessary to compensate for the thermal growth of the coupling shaft and the gearbox input pinion shaft when the unit reaches full speed operating temperature.

Figure 19.8 shows the coupling shaft arrangement. The axial thrust bearing, located on the gearbox input pinion shaft, carries the combined axial force (in a forward direction towards the GT) from the single helical gear train end thrust and that resulting from the cold prestretch of the high-speed coupling shaft. In the hot condition, at full operating power (steady state condition) an axial tension must be maintained in the drive train. This is because the output shaft bearing is designed to run with a tension loading if it is unloaded, or a compressive (aft) load is applied, the bearing will have a very short life.

During run-up of the GT the drive train only spends a very short time in the cold, full speed (7000 rpm) condition. The shafts quickly heat up to 120C. Because of centrifugal effects on the flexible diaphragm elements in the high-speed coupling shaft the tension force in the coupling varies with rotational speed. The relevant characteristics, based on type-test data for similar couplings, are shown in Fig. 19.8. The axial prestretch of the coupling can be adjusted by moving the GT forward or aft. The mounting system comprises a series of high-tensile linkages arranged as a three-dimensional spaceframe. During any axial adjustment of the GT position its lateral alignment, relative to the forward/aft axis of the vessel, must be maintained, otherwise the output bearing and gearbox input bearing may fail due to excessive lateral forces and vibration.

In passenger cars, trucks and tractors, a large proportion of closed-die hot forged steel components such as shafts, pinions, gears etc. are primarily used for strength, toughness, fatigue, along with dimensional accuracy. The paper provides an insight into the mechanism of grain growth in austenite during high temperature hot forging; stability of various precipitates in micro alloyed steels and its influence in restricting grain size; possible phase transformations during cooling; strengthening mechanisms and related aspects. Development of various forging grades over the years is analysed; it is shown that although the medium carbon, high manganese grades are predominantly used today in the hardened-and-tempered condition, several alternative grades are possible where the costly heat treatment and the accompanying risks of distortion / cracking could be avoided through air-cooling and yet obtain comparable or better properties. The requirement of through-thickness properties of large forgings and the role of acicular ferritic structures in very low carbon micro alloyed steel has been discussed. Demands such as machinabiity, weldability and the appropriate microstructures for the same have been looked into. Finally, it is shown that newer, non-steel materials (e.g. composites) as well as new versions of austempered ductile iron (ADI) are making their entry into the scenario which was so long considered a strong territory for steel forgings.

As in ball mills, the power draft of a rod mill is the product of capacity and work index, which is the energy required to break a mineral of a given size to the required size. The mill power is also increased by increasing the rod charge and the mill speed, while the mill power and capacity are both increased with increasing mill length.

For small ball mills, the power draw under dry batch grinding conditions was derived by Austin etal. [5] and the same considerations apply for rod mills. Equation (8.6) indicates that like any tubular mill the variation of mill power with speed in a rod mill is almost linear. This is true at the initial stages but breaks down when the critical speed is reached. At speeds in excess of the critical speed the power requirements decrease sharply. This is to be expected as rotation in excess of the critical speed results in the charge adhering to the inside liner and does not either cascade or cataract. Typical power requirements for two different mill loadings obtained in a laboratory size mill (0.6m 0.31m) with 20 lifters 25mm high are plotted in Figure8.8.

Figure8.8 shows the general characteristics of the change of mill power with mill speed for 17% and 40% mill loadings of a tumbling mill whose critical speed was 101rpm. It can be seen that at 40% loading the maximum mill power occurred at about 70% of the critical speed, while at a lower loading the maximum power drawn was nearly at the critical speed.

Equation (8.6) indicates that the power required is a function of the critical speed. Some manufacturers recommend an optimum speed of operation of their rod mills. For example, Marcy mills suggests that for their mills the peripheral speed should be governed by the relation

In industrial situations where conditions differ from Bonds set-up [4], Rowland and Kjos suggested in a series of papers [2,8] that Bonds Equation (3.25) can be used after correcting for different conditions encountered in industrial practice. Austin etal. [5] pointed out similar corrections required to Bonds equation to meet industrial conditions. These corrections are summarised below for specific conditions and are applicable to both rod and ball mills. More than one correction factor may be applicable. All factors are considered separately and the total correction is determined.

F2 is known as the inefficiency factor for the wet closed circuit grinding. It is a function of the sieve size used to determine the value of work index, Wi, and the percentage passing this control size. This function has been determined for different percentages passing the controlling sieve size and is shown in Figure8.9.

The value of Wi is best taken from an impact test or a rod mill grindability test, whichever is the greater. For a ball mill, the value of the constant in Equation (8.12) equals 4000 according to Rowland and Kjos [2].

For efficient operation of rod mills the feed should preferably be uniform in top size. The manner of feeding either from conveyors or directly from bins and chutes affects the power consumption and mill performance. To correct for feed preparation is difficult. The rule of thumb suggested by Rowland and Kjos [2] is summarised in Table8.2.

The above considerations for determining the mill power draw serve as a guide for the selection of mills for a particular job. Examples 8.1 and 8.2 illustrate the method of calculating mill power draw and also to compute the size of mill required for specific purpose.

A uniform discharge from a closed circuit jaw crusher is 200t/h. The crusher feeds a wet rod mill such that 80% of the crusher product passes a 16mm screen. The rod mill feeds a wet ball mill at a feed size of 1.0mm (1000 m) and produces a product with 80% passing a 150 m screen. The rod mill is in an open grinding circuit. Determine:1.the shaft power of the rod mill,2.the size of the industrial mill.

1.Correction factor F1 is not applicable.2.Correction factor F2 does not apply to rod mills.3.Correction factor F3 has to be considered after L and D are determined (usually towards the end of the computation). Hence, F3 will be determined later.4.Since the feed size is 16,000 m, correction factor F4 has to be determined.Use Equation (8.12) to determine FOPT, but first determine reduction ratio R.Reduction ratio (R)=16,000/1000=16.0Optimum feed size (FOPT ) = 16,000/(13 1.1/Wi)0.5=16,000 (13 1.1/13.5)0.5= 17,952 mSince the feed is less than the optimum, no correction is necessary.5.Correction factor F5 is not applicable.6.Correction factor F6 is applicable when R is between R*+2 and R*2.R* is estimated after mill size determination.7.Correction factor F7 is not applicable for rod mills.8.Correction factor F8 is not applicable as the circuit is a rod-ball mill circuit and the rod mill is fed from closed circuit crushing.

Preliminary selection of a commercially available rod mill may now be made from the manufacturers catalogue. For example, Allis Chalmers catalogue shows that the nearest mill size would require 655kW [2]. Such a mill would have the following tentative dimensions:Mill length=4.88mMill diameter=3.51m (inside diameter=3.31m)Rod length=4.72Rod load=40%Rod charge=90.7t

The diameter efficiency factor, F3 in step 2 can now be determined using Equation (8.10) (Figure8.10). As the ID of the mill has been provisionally established as 3.31m, thenF3=2.443.310.2=0.941 forD<3.81m

Referring again to Allis Chalmers rod mill performance in Appendix B-4, for a mill power draw of 618kW, the following mill will be finally suitable:Mill length=4.86mMill diameter=3.35m (inside diameter=3.15m)Rod length=4.72mRod load=45%Rod mass=93.5t

To determine the commercially available ball mill that would suit the conditions, refer to the manufacturers literature, for example, Allis Chalmers as published by Rowland and Kjos [2]. From the tabulated data, the mill charge and other characteristics corresponding to a power draw of 1273kW areBall mill length=4.57mBall mill diameter=4.57m (inside liner diameter=4.39m)Ball mill load=35%Ball charge mass=113tBall size=64mm

Referring again to Allis Chalmers ball mill performance table, for a mill power draw of 1412kW, the same mill dimensions will be suitable if the ball charge is increased to 45% with a charge mass of 144t.

Since between-bearing helical gears are the most common type on site, only this type will be covered. However, the relations discussed, with minor modifications, will also apply to internal and external spur gears. Figure 4.4.5 shows the reaction forces that act on a helical pinion tooth.

The axial load, WA, is calculated directly as shown in Figure 4.4.5, and can be applied to either the gear shaft, pinion shaft, or divided between the gear and pinion shaft. Most gear designs absorb all thrust on the gear shaft (low speed), since this usually results in the lowest thrust bearing losses.

Shaft centerline position is monitored by two proximity probes that are mounted at each radial bearing. They will record the position of the shaft and therefore the load angle of the shaft in the bearing. The load angle in gear applications changes with transmitted torque (power/speed).

Critical integral gear compressors have required emergency shutdown due to excessive radial bearing wear that did not cause high levels of vibration prior to shutdown (high vibration occurred when Babbitt material was excessively worn).

Heavily loaded radial bearings do not exhibit high vibration, and can go undetected if shaft position is not monitored, and bearing pad temperature probes are not located at the load point of the bearing.

This best practice has been used since the 1980s to optimize highly loaded gear radial bearing life, by predicting and recommending machine shutdowns at convenient times, thus eliminating costly emergency shutdowns.

Tumbling mills are most commonly rotated by a pinion meshing with a girth ring bolted to one end of the machine (Figure 7.10). The pinion shaft is either coupled directly or via a clutch to the output shaft of a slow-speed synchronous motor, or to the output shaft of a motor-driven helical or double helical gear reducer. In some mills, electrical thyristors and/or DC motors are used to give variable speed control. Very large mills driven by girth gears require two motors, each driving separate pinions, with a complex load sharing system balancing the torque generated by the two motors. (See also Knecht, 1990.)

The larger the mill, the greater are the stresses between the shells and heads and the trunnions and heads. In the early 1970s, maintenance problems related to the application of gear and pinion and large speed reducer drives on dry grinding cement mills of long length drove operators to seek an alternative drive design. As a result, a number of gearless drive (ring motor) cement mills were installed and the technology became relatively common in the European cement industry.

The gearless drive design features motor rotor elements bolted to a mill shell, a stationary stator assembly surrounding the rotor elements, and electronics converting the incoming current from 50/60Hz to about 1Hz. The mill shell actually becomes the rotating element of a large low speed synchronous motor. Mill speed is varied by changing the frequency of the current to the motor, allowing adjustments to the mill power draw as ore grindability changes.

The gearless drive design was not applied to the mills in the mineral industry until 1981 when the then-worlds largest ball mill, 6.5m diameter and 9.65m long driven by a 8.1MW motor, was installed at Sydvaranger in Norway (Meintrup and Kleiner, 1982). A gearless drive SAG mill, 12m diameter and 6.1m length (belly inside liners) with a motor power of more than 20MW, went into operation at Newcrest Minings Cadia Hill gold and copper mine in Australia, with a throughput of over 2,000t h1 (Dunne et al., 2001). Motor designs capable of 35MW have been reported (van de Vijfeijken, 2010).

The major advantages of the gearless drive include: variable speed capacity, removal of limits of design power, high drive efficiency, low maintenance requirements, and less floor space for installation.

Packaged Centrifugal Compressors are integrally geared, skid-mounted centrifugal units consisting of a driver (motor or turbine), a bull gear, and up to four pinions. Each pinion shaft terminates in an impeller. These machines are often described as packaged air machines, since the most common service is general plant compressed air supply. Generally, the driver and bull gear speed is 3,600 rpm or less, and the pinion speeds can be as high as 60,000 rpm. These machines are produced as a package with the entire machine mounted on a common foundation which also includes a panel with control and monitoring instrumentation. Because of the large number of these machines manufactured, proper monitoring locations for proximity probes have been established by the various machine manufacturers. Nearly all of these machines are supplied by the OEM with one proximity probe per impeller (one or two per pinion) and sometimes one probe on the bull gear shaft. In some instances, manufacturers have been able to respond to user specifications by supplying X-Y proximity probes and Keyphasor probes on the various shafts.

Process Centrifugal Compressors are usually larger than packaged units and have radial-flow or axial-flow stages usually mounted in the center span of the rotor between two fluid film bearings. These machines should be monitored with XY radial proximity probes at each journal bearing and at least one axial probe at the thrust bearing. If thrust position measurement is connected into automatic shutdown, then two axial probes should be installed in a voting logic configuration to reduce the possibility of false trips. If radial vibration channels are also wired into automatic shutdown, then extra false trip protection should be incorporated in the radial vibration monitors.

It is not necessary to mount a Keyphasor probe on the compressor if the coupled driver operates at the same speed and is equipped with a Keyphasor. However, if a gearbox is part of the system, then the compressor(s) on the opposite side of the gearbox from the driver should have a Keyphasor. Temperatures of all bearings, oil supply, and ambient conditions should also be monitored. On axial compressors, an accelerometer can be a useful auxiliary measurement for determining blade disturbances.

Reciprocating Compressors of the horizontal type can usually be monitored by using X-Y proximity probes to observe the piston rod or plunger of the compressor in addition to X-Y probes at accessible main crankshaft bearings. Monitoring the average position of the plunger can determine rod misalignment, packing wear, and wear on the sliding element and cylinder liner. Monitoring the dynamic motion signal of the probe can determine rod vibration or flexure (deflection). This is most necessary on hypercompressors. A Keyphasor could be installed on one of the drive rods to provide speed and timing information. An alternative location for the Keyphasor may be found on the crankshaft or motor driver. The Keyphasor should be installed so that the voltage pulse occurs when one cylinder is at top dead center to observe the relationship of all plunger positions as a function of stroke.

Screw Compressors have two rotors with interlocking lobes and act as positive displacement compressors. If the machine is equipped with fluid film bearings, the optimum installation should include X-Y proximity probes at each radial bearing of each rotor as well as at least one axial probe for thrust position measurement of each rotor. Thrust position monitoring of each rotor is very important because a thrust bearing failure means an axial rub will occur between the rotating elements. If rolling element bearings are used, monitoring with casing velocity transducers and Dual Path Monitors is acceptable; however, due to the very close axial clearances between rotors, axial shaft position monitoring is still of importance. The case-mounted transducer should be a velocity pickup with a Dual Path Monitor measuring both velocity and displacement vibration. Temperatures of machine components should also be monitored. Two Keyphasor probes, one for each rotor, would be desirable.

The location of the gear-box thrust bearing will determine whether thrust transmission will occur for a given machinery train arrangement. For the train configuration shown in Figure 3-77, the thrust bearing is generally part of the high-speed pinion shaft assembly. Using a coupling friction coefficient of 0.3, the pinion thrust bearing could be exposed to a maximum axial thrust of

However, the maximum possible axial force imposed on the gear mesh on Figure 3-77 is not related to F1, but rather, as will be seen later, to the tangential driving load FT and the magnetic centering force MA of the motor.

If the thrust bearing is mounted on the low-speed gear shaft, as shown in Figure 3-78, it could be exposed to the same maximum thrust as the bearing shown in Figure 3-77. However, the thrust force is now transmitted across the gear mesh. The maximum possible axial force imposed on the mesh is now not only a function of FT but of F1 as well.

The analysis can, of course, be extended to drivers other than motors. Unlike motors, which are generally furnished with axially free-floating rotors, these other drivers will probably incorporate thrust bearings.

There are several designs available, operated by a single acting cylinder with spring return, a double acting cylinder or twin cylinders. The basic principal is the same in each: the piston rod of a pneumatic cylinder becomes a rack which rotates the pinion shaft. In a twin cylinder arrangement it is possible to have a three- or four-position actuator according to which ports are pressurised, as shown in Figure 10. The angular rotation is limited in principal only by the stroke length of the cylinder, but in practice, standard units rarely exceed about 360. Up to 400 Nm is possible with this type of design.

The turbine shaft serves the function of a pinion shaft of the gearbox, thereby increasing the compactness and sturdiness of the design. This is especially suited as a cost-efficient solution in ORC-based power plant applications where a single turbine alone cannot transfer the entire power provided by the process. The same applies to processes with multiple pressure levels.

The conclusion of this is that the inward radial flow turbine design form Atlas Copco is extremely reliable and requires little maintenance, even if this is a high performance design with advanced features.

This high reliability and low maintenance has been verified in hundreds of turbines which have been made by Atlas Copco and delivered not only to ORC plants, but also to applications with similar (and not quite so similar) fluids such as natural gas condensate removal, liquid natural and petroleum gas, air separation, and so on.

electrolytic starter (lrs) for high power slipring motors - epm - electrolytic starters for slipring motors aoip

electrolytic starter (lrs) for high power slipring motors - epm - electrolytic starters for slipring motors aoip

EPM electrolytic starter can be used to start slipring motors from 500 kW to 20,000 kW. It will supply the power necessary to drive the motor by resistance variation. Designed for controlled starting and speed control of large slipring motors in demanding applications, the EPM liquir rotor starters ensure a smooth progressive acceleration of installations such as:

They are widely used in various industries such as mines, quarries, cement plants, water treatment and associated industries. They are also adapted to specific applications such as car fragmentisers, plastic mixers and sugar cane knives.

Several models and options are available according to the starting power required, the inertia of the driven machine and the application. Feel free to contact us and explain us about your requirements (Use the section Necessary information to quote a starter), our staff will quote the suitable starter and the eventual necessary options.

Tank: The tank is manufactured with heavy duty sheet steel 30/10 to 50/10 mm gauge, and is normally supplied complete with lifting eye bolts Tank capacity and dimensions are determined by the motor rating.

The moving electrodes travel vertically inside the insulating container, guided by a nylon rod. The assembly is supported by two brass rods fixed to a transversal carrier which is common to all three electrodes and constitutes the neutral point.

Control and interlocking: Limit switches are incorporated to control the geared motor, and to power the shorting contactor which shorts out the residual resistance at the end of the run up time. The geared motor is fitted with

This type of starter is designed to provide the optimum starting characteristic, which results in smooth progressive acceleration to full load speed. It can also be used for speed variation and torque control. Plug braking can also be implemented with this system.

AOIP starters for slipring motors, also named electrolytic starters or Liquid Resistance Starters (LRS), use mobile electrolytic resistances. Each starter is made of 3 tanks (one per phase) filled with conductive liquid named electrolyte (water mixed with salt, usually Sodium Carbonate) and two immerged electrodes. One is fixed to the bottom of the tank, the other travels vertically inside the insulating container, guided by a nylon rod. The electrodes are in stainless steel (or bronze in option), shaped in concentric cylinders which merge with each other in the minimum resistance position.

The resistance value depends on the distance between the electrodes, Sodium Carbonate concentration, and electrolyte temperature. The level and temperature of the electrolyte are controlled by a float and thermostats triggering an alarm when set limits are reached.

By distance variation between the electrodes, we get an accurate variation of resistance, thus a supply voltage adjustment and a reduction of starting current and torque, purpose of the system. (see Why using a starter for slip ring motors)

The starter slowly decreases the resistance, ensuring a progressive starting of the driven machine, unlike step starting due to starters with fixed electrodes. At the end of the starting process, the resistance is short-circuited.

Some LRS manufacturers do not use mobile electrodes but prefer to vary the level of electrolyte to modify the resistance value. This solution implies using circulation pump which has two disadvantages:

In order to start a slip ring motor, the accelerating torque has to be sufficiently higher than the resisting torque. Higher is the inertia, higher has to be the accelerating torque, otherwise the start will be very long and the motor can be damaged by the heating.

In order to reduce the starting current and starting torque, it is necessary to decrease the supply voltage of the motor. If the supply voltage is divided by two, the absorbed current is divided by the same ratio of two.

For example, for a motor with a ratio starting torque / nominal torque of two and a ratio starting current / nominal current of six under a nominal voltage of 400 V, if the supply voltage is reduced to 200 V the absorbed current is only three times the nominal current. Furthermore, the torque is reduced by the square of the supply voltage reduction, by reducing the supply voltage from 400 V to 200 V, the motor torque is 2/(2x2) ie 0.5 times the nominal torque.

Different models exist to suit the power of the starter(s) and the inertia of the driven machine.The ranges of EPM starters given above are theoretical only, as ranges will depend on many further non negligible parameters such as starting conditions, starting time and cadenza, torque, type and load of the driven machine, ambient temperatureRotor voltage between rings: 3,500 V maxStandard starting times: 20, 30, 40, 60, 80, 130 s factory presetLevel of electrolyte monitored by magnetic floating systemElectrolyte temperature monitored by thermostatsElectrolyte cooling down by natural convection and agitator mixingLow current density of electrodes: about 1 A/cm.

AOIP offers a complete range of services: technical assistance and advice, calibration and verification by the SOFIA laboratory, equipment repair, maintenance, assistance and metrological monitoring, equipment rental, training, installation and commissioning of your equipment. More informations

electromechanical dynamic behaviour and start-up evaluation of tumbling ball mills

electromechanical dynamic behaviour and start-up evaluation of tumbling ball mills

Weidong Lv, Guoqiang Wang, He Tian, "Electromechanical Dynamic Behaviour and Start-Up Evaluation of Tumbling Ball Mills", Mathematical Problems in Engineering, vol. 2018, Article ID 3515308, 13 pages, 2018. https://doi.org/10.1155/2018/3515308

This paper presents a dynamic simulator of the electromechanical coupling start-up of a ball mill. The electromechanical coupling model based on the dynamic model of the ball mill, the characteristic equation of the clutch, and the dynamic model of the induction motor is established. Comparison between the simulation results of angular speed, load torque and current obtained from the model, and the experimental results is conducted to validate the correctness of these simulation results. Results show that the simulation results of the electromechanical model are highly consistent with the experimental results. Two indexes are proposed for evaluation. Finally, a 4500kW ball mill is used to analyse the start-up process with different operation parameters of the air clutch. The effect of the engagement time and the pressure of the air clutch on the torque, current, and shock extent is analysed. Moreover, the optimum inflation time is determined.

Ball mill is an important equipment in the field of mineral size reduction. Unlike semiautogenous (SAG) mills that operate with adjustable frequency drives, the ball mills work at a fixed speed [1]. Current research mainly focuses on the power draw model. However, the shutdown maintenance and the component consumption caused by the failure of the clutch friction disc wear and motor overheating, which increase the operating costs, are also important research topics. Therefore, the mill should be studied from the perspective of electromechanical coupling rather than only considering the mechanical system of barrel material. Castro and Valenzuela [2, 3] studied the electromechanical coupling model of a SAG mill. The simulation results were compared with the field record data during start-up and shutdown of the SAG mill. Guerrero and Pontt [4] investigated the oscillatory torque caused by dead time in the current control of a high-power gearless mill drive. Szolc et al. [5] examined the interaction of the electromechanical coupling of an asynchronous motor drive system. Laylabadi and Symonds [6] analysed the control system of an 18 MW ball mill drive system.

The establishment and validation of the tumbling mill models have been widely conducted. Morrell [7] set up a power draw model of SAG mill that considers the slurry phase. Zolghadri et al. [8] developed a novel theory of breakage function. The method is considered for energy in comminution modelling of specific breakage energy. Salazar et al. [9, 10] established a dynamic model of SAG mill based on feed and discharge and the model predictive control method. Djordjevic et al. [11] used a prototype to study the law of particle breakage and energy distribution in the mill. A 600mm tumbling mill prototype was used to verify the DEM model in [12]. Therefore, the ball mill model is typically verified with a reduced proportion of prototypes.

The drive motor of the large ball mill is usually required to start with a reduced-voltage method [13]. Therefore, the large ball mill usually uses air clutch to assist start-up. During the start-up process, the motor is started up to the rated speed. Then, the ball mill is started through the engagement of the clutch. This study focuses on the dynamic behaviours of a ball mill and the dynamic characteristics of the induction motor. The electromechanical coupling model is established, and simulation of a 4500kW ball mill during start-up is conducted.

Figure 1 shows a ball mill driveline. The main shaft of the drive motor is connected to the driving disc of the air clutch. The driven disc of the air clutch is connected to the pinion shaft. is the moment of inertia of the motor rotor and clutch driving disc. is the equivalent moment of inertia of the pinion shaft. is the reduction ratio. is the moment of inertia of the rotary sections and materials of the ball mill.

According to Figure 1, the dynamic model of the ball mill iswhere is the rotation resistance coefficient of the drive motor, is the friction torque, and the rotation direction of the ball mill is determined by the sign function sign().

The large ball mill commonly uses a double-row heavy-duty air clutch. The pressure of the friction surface is determined by the structure of the clutch. For the radial air clutch, the radial pressure generated by the compressed air can be calculated by the following formula:where is the minimum radius of the air tube, is the width of the air tube, is the minimum working pressure, and is the initial pressure.

After the air tube is inflated, it expands in the radial direction and overcomes the force of the return spring. As a result, the friction plate can hold the hub. Therefore, the pressing force iswhere is the force of the return spring, is the sum of the radial pressure generated by the compressed air through the air tube, and is the sum of the centrifugal forces generated by all the parts that undergo radial displacement during operation of the clutch.

Figure 2 shows the curve of the pressure versus inflation time in accordance with the mechanical design manual. In 00.528, the airspeed in the air tube is the speed of sound, and the inflator flow of the tire is constant. When the time exceeds 0.528, the pressure difference between the air tube and air tank decreases, thereby decreasing the airspeed to the subsonic. The time constant can be calculated as follows:where is the container volume, s is the effective cross-sectional area of the inflation valve, is the isentropic exponent, and =1.3 for air.

The clutch inflation time is generally around 10 s, and the maximum friction torque of the clutch is generally greater than the maximum load torque of the ball mill. Therefore, this study assumes that the air pressure characteristics of the clutch are as follows:where is a constant related to the inflation speed and is the air tank pressure.

The time from the start of the inflation to the contact (but not sliding) of the driving and driven discs is generally less than 3 s. The inflation time is usually from 6 s to 12 s. The relationship between clutch friction torque and pressing force can be expressed aswhere is the coefficient of frictional torque and is the surface area of friction.

The radius and angle position of the mill charge centre of mass is commonly used for the ball mill dynamic model during start-up. This model considers that the mill load maintains its resting shape and that the mill is lifted to a certain angle depending on the mill speed and existing operating conditions during its rotation [3].

As shown in Figure 3, the mill torque is obtained aswhere Mload is the mass of charge in the barrel, is the distance between mill centre and the centroid of the charge, and is the deflection angle of the mill [14].

According to formulas (9), (10), and (11), the state equation of the induction motor can be obtained as follows:where is the number of pole pairs, the electromagnetic time constant of the rotor is =/, =1-/(LsLr) is the leakage coefficient of the motor.

The evaluation of the start-up process of the ball mill aims to ensure successful start-up without the drive or clutch protection device that causes the mill to stop. Considering this condition, the start-up process is made as short as possible to reduce friction disc wear and heat generation. Therefore, the quantitative evaluation indexes of the start-up process of the ball mill can be set as the shock extent and motor protection to avoid vibration damage and tripping of the power.

(1) Shock Extent. During start-up of the ball mill, if the clutch engagement time is too fast, then the impact will cause damage to the drive motor and the conventional system. The shock extent can be expressed as the rate of change in the spindle angular acceleration [15], which is

(2) Motor Protection. Too fast clutch engagement may cause the motor to the locked rotor and the relay to act. Therefore, during start-up, the drive motor should not be locked rotor and no overcurrent protection should be triggered. The current limit must be set in accordance with the motor parameters.

The simulation is based on a large ball mill with the parameters shown in Table 1. The motor that matches the ball mill is a 4500kW induction motor with the parameters shown in Table 2. The parameters of the matching air clutch are shown in Table 3.

The inflation time range of the air clutch is generally from 6 s to 12 s. In accordance with the actual situation, this section sets the inflation time () of the air clutch to =6 s, =7 s, =8 s, =9 s, and =10 s. The maximum friction torque is 100% of the maximum friction torque of the air clutch. Consistent with the actual start-up procedure, the results are plotted from the moment the clutch starts to engage for ease of analysis.

Figure 4 shows the speed comparison of the start-up process for different inflation times. Figure 4(a) shows the speed curve when the inflation time is 8 s. As the rotational speed of the driven disc increases, the motor speed decreases. After the speeds of the driving disc and the driven disc are synchronized, the ball mill accelerates to the working speed. Because the mechanical characteristics of the large induction motor are hard, the amplitude of the speed decrease is small. Figure 4(b) shows the comparison of the angular velocity of the driven disc. The start-up time increases along with the increase in the inflation time. When =6 s, the start-up time is 5.6 s and the engagement time is 5.2 s. When =7 s, the start-up time is 6.2 s and the engagement time is 5.9 s. When =8 s, the start-up time is 7.0 s and the sliding friction time is 6.7 s. When =9 s, the start-up time is 7.9 s and the engagement time is 7.6 s. When =10 s, the start-up time is 8.8 s and the engagement time is 8.5 s. In general, as the inflation time decreases, the friction time of the driving and driven discs increases and the start-up time of the ball mill is slightly shorter than the inflation time. In the five cases, the ball mill barrel begins to rotate at approximately 1 s and the speed gradually increases. As the materials are lifted, the load torque gradually increases. The friction torque of the clutch is less than the load torque, and the rotational speed of the barrel decreases. The smaller the friction torque, the more the speed drops. As the magnitudes of the friction torque increase, the rotational speed increases again. The rotational speeds have been increased until the speeds of the driving and driven discs are the same for =6 s, =7 s, and =8 s, respectively. The magnitudes of the rotational speeds fall again after the second increase (=9 s and =10 s). As the friction torque increases, the rotational speed increases again, which ends the friction process.

Figure 5 is the angle curve of the barrel. As shown in the picture, for all 5 inflation time, the angle of barrel increases steadily. Before the barrel was rotated by 90, the rotation angle in the 5 inflation times shows changes in the growth rate. This law is consistent with the change law of the angular speeds in Figure 4 and is related to the torque.

Figure 6 shows the change in electromagnetic torque with the variation in inflation times. All the magnitudes of the electromagnetic torque increase as the friction torque increases during the friction process, and these magnitudes decrease as the load torque of the motor decreases when the rotational speeds of the driving and driven discs are synchronised. At =6 s, =7 s, and =8 s, the electromagnetic torque is less than the load torque of the ball mill during the friction process. This variation is shown in Figure 7. This condition corresponds to the decrease in the rotational speed in Figure 4. Similarly, the electromagnetic torque is two times less than the load at =9 s and =10 s. Therefore, the rising rate of the electromagnetic torque is positively correlated with the inflation time. Moreover, the electromagnetic torque at =6 s is slightly larger than that at the other conditions.

Figure 7 shows the curve of the theoretical friction torque of the clutch and the load torque of the ball mill. As the model of the mechanical system, the theoretical friction torque and the load torque are determined according to (7) and (8), respectively. For each of the 5 inflation times, the friction torque is less than the load torque of the ball mill for several times during the clutch engagement. This is shown in Figure 4 and Figure 5 as the decrease in angular speed and the increase in the deceleration of rotation angle. However, because the angular speed does not drop to 0 and the barrel does not reverse, the ball mill is successfully start-up. The variation of electromagnetic torque can also be observed from Figure 6. The electromagnetic torque changes based on the load torque of the motor. At the beginning of the air clutch engagement, the load torque of the motor is equal to the friction torque. When the friction torque is greater than the initial load torque (the static friction torque and the inertia torque), the driven disc beginning accelerates. When the friction torque is less than the load torque of the ball mill, the angular speed of the driven disc decreases. After the speeds of the driving disc and the driven disc are synchronized, the load torque of the motor is equal to the load torque of the ball mill. Although the theoretical friction torque still increases, as shown by the solid line in the Figure 7, the friction torque that the air clutch transmission is equal to the load torque of the ball mill. In addition, with the rotational speed and angle of the barrel increase, the material begins to cascade and fall, which causes the load torque of the ball mill decrease. Therefore, after the speeds of the driving and driven disc synchronized, the friction torque rapidly drops with the decrease of the load torque. In Figure 6, the data shows that the electromagnetic torque is linear increased at the beginning and then decreases.

Figure 8 shows the line current amplitude during the ball mill start-up. The currents which consist with the change in the electromagnetic torque in Figure 6 are gradually increased from the no-load current to the maximum value, dropped and stabilised at working current. The working current is 349.2 A. The maximum values of the phase current at =6 s, =7 s, =8 s, =9 s, and =10 s are 572.6, 555.4, 552.5, 557.7, and 557.4 A, respectively. Therefore, the suitable inflation time is from 7 s to 10 s.

Figure 9 shows the curve of the shock extent. The shock extent is significant only in the friction process because the difference in the speeds of the driving and driven discs leads to impact. Thus, the peaks in the synchronous process can only reflect the value of acceleration. In general, the maximum peak of the shock extent decreases as the inflation time increases. The number of peaks increases as the start-up time increases. At =6 s, =7 s, and =8 s, three peaks are observed during the mill start-up. Four peaks are observed at =9 s and =10 s. The maximum values of shock extent are 70.4 (=6 s), 60.3 (=7 s), 52.8 (=8 s), 46.9 (=9 s), and 42.2 rad/s3 (=10 s). The optimal inflation time should be between 8 s and 9 s with the presence of shock extent.

The simulation results show that =8 s is appropriate for the ball mill. It not only ensures the success rate of the start-up but also will not cause great impact and shock on the drive motor, the transmission system, and the mill.

(1) Too Short Inflation Time. =3 s and =4 s are used in this simulation. Figure 10 shows the angular speed curves during the start-up process. In general, the motor can drive the ball mill to start. When =3 s, the clutch reaches the synchronous process and working speeds at 2.5 and 2.9 s, respectively. When =4 s, the clutch reaches the synchronous and working speeds at 3.6 and 3.9 s, respectively.

Figure 11 shows the curve of the rotation angle of the barrel. The ball mill is successfully start-up in both inflation time. For =3 s, the rotation angle increases smoothly. For =4 s, the increase in rotation angle becomes slower at 2.8 s because of the angular speed decrease. After the clutch is synchronized, the rotation angle constant increases in the working speed.

Figure 12 shows the curve of the electromagnetic torque versus time. The electromagnetic torque increases with the increase in the theoretical friction torque of the clutch. The mill starts to rotate at 0.5 s. After the synchronisation of the driving and driven discs, the electromagnetic torque decreases to equal to the load torque.

Figure 13 shows the curves of the theoretical friction torque and the load torque. The theoretical torque increases linearly from 0, and then reaches its maximum at the end of inflation time. For =3 s, the theoretical friction torque is always greater than the load torque. Therefore, the angular speed of the driven disc does not decrease in Figure 10. For =4 s, the theoretical friction torque is less than the load torque once. This is corresponding to the decrease in the angular speed of the driven disc and the reduction of the rotation angle growth rate in Figures 10 and 11, respectively.

Figure 14 shows the curve of the current amplitude during the start-up process. The maximum currents are 549.3 A (=3 s) and 588.4 A (=4 s). For =4 s, the maximum current is 1.6 times more than the operating current. Although the peak is much smaller than the current during the motor start-up, it is still greater than =78 s.

Figure 15 shows the curve of the shock extent. The peak of the shock extent during the friction process is 141 rad/s3 (=3 s). The shock extent reaches its peak for 1 s and afterward drops. The first peak of the shock extent for =4 s is 105.7 rad/s3 at 1.1 s. The second peak is 41 rad/s3 at 2.3 s. The shock extent is twice as large as =78 s. Therefore, the ball mill will be damaged during an actual start-up.

(2) Too Long Inflation Time. =12 s, =15 s, and =18 s are used to simulate too long inflation time. Figure 16 shows the angular speed curves of the driven disc. The result of =12 s is same as that of =10 s. In particular, the speed curve has two fluctuations in the friction process. When =15 s and =18 s, the curve has three fluctuations. The duration of the friction process increases with the increase in the inflation time (9.8 s at =12 s, 12.3 s at =15 s and 14.4 s at =18 s). The start-up of the ball mill is close to 20 s.

Figure 17 shows the curve of the barrel rotation angle when the inflation time is too long. The rotation angle growth rate is consistent with the variation of the angular speed. For the three inflation time, although the angular speed was very small at 4.2 s, the barrel did not reverse. Therefore, the ball mill can be start-up successfully. Also, in the clutch engagement, the longer the inflation time the slower the increase in the angle of the barrel. After the speeds of the air clutch are synchronized, the acceleration is constant that the rotational speed was reached the working speed.

Figure 18 shows the curve of the theoretical friction torque and the load torque. For =12 s, the theoretical friction torque is less than the load torque twice in the friction process, which is the same with the variation of the angular speed. For =15 s and 18 s, the theoretical friction torque is less than the load torque for three times in the friction process. After the angular speeds synchronized, the actual friction torque is equal to the load torque, and decrease to the rated working torque of the ball mill. This variation is shown in Figure 18 as the process of current decrease.

Figure 19 shows the curves of the current versus time. The peaks of the current are 540 A (=12 s), 541 A (=15 s), and 527.7 A (=18 s). Too slow clutch engagement increases the duration of the peak of the current. Moreover, the peak currents of three are similar. All the maximum currents of three are smaller than the maximum friction torque. Therefore, the extreme load of the drive motor does not change considerably as the clutch engagement slows down. The slower the engagement is, the longer the duration of the high current continues. In the meantime, the pressure on the grid and the drive motor increases. Thus, this condition should be avoided.

Figure 20 shows the shock extent. When =12 s, the duration is 6.8 s and the peak is 35.2 rad/s3. When =15 s, the duration is 8.1 s and the peak is 28.2 rad/s3. When =18 s, the duration is 10.1 s and the peak is 23.5 rad/s3. Compared with those in =78 s, the amplitude is reduced by half, but the duration is doubled, which is inappropriate.

The results of this section show that the inflation time significantly affects the heat generation of the friction discs and the duration of the current. The change in the operation parameters of the clutch should be strictly avoided after the optimal inflation time of the air clutch is determined.

This section assumes that the pressure of the air tank has changed. Therefore, the maximum friction torque () of the clutch decreases or increases. The settings used are as follows: =7 s and =80%, =, =120%, =160%, and =200%.

Figure 21 shows the curve of the angular speed of the five conditions. The ball mill can successfully start even under air pressure loss (=80%). In the meantime, the duration of the high current and friction process increases. In addition, if <80%, then the ball mill will fail to start. The greater the friction torque at constant , the faster the clutch engagement and the smaller the fluctuation of speed. The speed has two fluctuations at =80%. The others only have one fluctuation. The fluctuation decreases with the increase in the friction torque.

Figure 22 shows the rotation angle curve of the ball mill start-up with different air pressures of the air clutch. The increasing rate of the rotation angle corresponds to the change in the angular speed. For the five , although the angular speeds are all decreased during the mill start-up, the barrel is not reversed. Therefore, the ball mill can be successfully started under the five working conditions. Moreover, the smaller the air pressure is, the slower the rotation angle increasing rate will be.

Figure 23 shows the curve of the theoretical friction torque and the load torque. When the inflation time is constant, increasing the air pressure is equivalent to increasing the inflation speed. For =80%, the number of times which the theoretical friction torque is less than the load torque is more than other conditions, and the duration is longer. This results in a greater fluctuation in the angular speed, as shown in Figure 21. It is also reflected in the increase in the duration of the shock extent in Figure 25. Therefore, the air pressure loss must be avoided.

Figure 24 shows the electromagnetic torque during the start-up process. The friction torque of the clutch is sufficiently large. Thus, the increase in cannot affect the peak of the electromagnetic torque. The electromagnetic torque does not reach the friction torque during the friction process. Therefore, the increase in the air pressure corresponds to the decrease in the inflation time.

Figure 25 shows the shock extent of the five conditions. As decreases, the amplitude of the shock extent decreases and the duration increases. When =80%, the peak of the shock extent is similar to =. Therefore, the air pressure should be constant after the clutch parameter is determined.

In this study, the electromechanical coupling dynamic model of a single-motor edge-driven ball mill is established. Then, the start-up process of a large-scale ball mill is simulated. The electromechanical performance of the ball mill is evaluated in terms of the shock extent and current. The conclusions based on the observations, calculations, and analyses are as follows.

Changes in the parameters of the air clutch can cause the changes in the clutch engagement. In the friction process, the angular speed of the driven disc may decrease during the increase of the synchronous rotational speed. The reason for the fluctuation can be explained as follows: at this moment, the friction torque that the air clutch can transmit is smaller than the torque generated by the rotation of the ball mill. However, as long as the barrel does not reverse and the rotation angle keep increasing, the angular speed of the driven disc fluctuation will not cause failure of the ball mill start-up.

The shock extent is too large when is too short, and this condition will damage the air clutch. Although the shock extent decreases with the increase in , the duration of the peak current and the friction time of the clutch increase, which generates a large amount of heat. As a result, the life of the drive motor and the clutch friction disc decrease, and this condition should be avoided.

When the air pressure of the clutch increases, which is similar to the decrease in , the shock extent increases. Although some pressures are lost, the ball mill can still start up (20% in this study). However, the duration of the current will increase. Therefore, the pressure should be monitored. If the pressure is lost, the start-up of the ball mill should be stopped.

This work was supported by the National Natural Science Foundation of China (Grant no. 51775225) and the Shanxi Province Coal Basic Key Technologies Research and Development Program (Grant no. MJ2014-02).

Copyright 2018 Weidong Lv et al. This is an open access article distributed under the Creative Commons Attribution License, which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.

ball-milled biochar for alternative carbon electrode | springerlink

ball-milled biochar for alternative carbon electrode | springerlink

Ball-milled biochars (BM-biochars) were produced through ball milling of pristine biochars derived from different biomass at three pyrolysis temperatures (300, 450, and 600C). The results of scanning electron microscopic (SEM), surface area, hydrodynamic diameter test, and Fourier transform infrared spectroscopy (FTIR) revealed that BM-biochars had smaller particle size (140250nm compared to 0.51mm for unmilled biochar), greater stability, and more oxygen-containing functional groups (2.24.4mmol/g compared to 0.82.9 for unmilled biochar) than the pristine biochars. With these changes, all the BM-biochar-modified glassy carbon electrodes (BM-biochar/GCEs) exhibited prominent electrochemical properties (e.g., Ep of 119254mV compared to 850mV for bare GCE). Cyclic voltammetry (CV) and electrochemical impedance spectra (EIS) show that ball-milled 600C biochar/GCE (BMBB600/GCE and BMBG600/GCE) had the smallest peak-to-peak separation (Ep=119 and 132mV, respectively), series resistance (RS=88.7 and 89.5, respectively), and charge transfer resistance (RCT=1224 and 1382, respectively), implying its best electrocatalytic activity for the reduction of Fe(CN)63. It is supposed that the special structure (i.e., internal surface area, pore volume, oxygen-containing functional groups, and graphitic structure) facilitates the electron transfer and reduces interface resistance. Economic cost of BM-biochar/GCE was 1.97107 USD/cm2, much lower than that of a low-cost platinum electrode (0.03 USD/cm2). The results indicate potential application of the novel BM-biochar for low cost and high efficient electrodes.

Chen L, Tang Y, Wang K, Liu C, Luo S (2011) Direct electrodeposition of reduced graphene oxide on glassy carbon electrode and its electrochemical application. Electrochem Commun 13:133137. https://doi.org/10.1016/j.elecom.2010.11.033

Chen J, Chen Q, Ma Q, Li Y, Zhu Z (2012) Chemical treatment of CNTs in acidic KMnO 4 solution and promoting effects on the corresponding PdPt/CNTs catalyst. J Mol Catal A:Chem 356:114120. https://doi.org/10.1016/j.molcata.2011.12.032

Chen Z, Xiao X, Chen B, Zhu L (2015) Quantification of chemical states, dissociation constants and contents of oxygen-containing groups on the surface of biochars produced at different temperatures. Environ Sci Technol 49:309317. https://doi.org/10.1021/es5043468

Choi J, Jin J, Jung IG, Kim JM, Kim HJ, Son SU (2011) SnSe2 nanoplate-graphene composites as anode materials for lithium ion batteries. Chem Commun (Camb) 47:52415243. https://doi.org/10.1039/c1cc10317b

Gao J, Wang W, Rondinone AJ, He F, Liang LY (2015) Degradation of trichloroethene with a novel ball milled Fe-C nanocomposite. J Hazard Mater 300:443450. https://doi.org/10.1016/j.jhazmat.2015.07.038

Gu Y, Wang B, He F, Bradley MJ, Tratnyek PG (2017) Mechanochemically sulfidated microscale zero valent iron: pathways, kinetics, mechanism, and efficiency of trichloroethylene dechlorination. Environ Sci Technol 51:1265312,662

Guai GH, Leiw MY, Ng CM, Li CM (2012) Sulfur-doped nickel oxide thin film as an alternative to Pt for dye-sensitized solar cell counter electrodes. Adv Energy Mater 2:334338. https://doi.org/10.1002/aenm.201100582

Huggins T, Wang HM, Kearns J, Jenkins P, Ren ZJ (2014) Biochar as a sustainable electrode material for electricity production in microbial fuel cells. Bioresour Technol 157:114119. https://doi.org/10.1016/j.biortech.2014.01.058

Ju MJ, Jeon IY, Lim K, Kim JC, Choi HJ, Choi IT, Eom YK, Kwon YJ, Ko J, Lee JJ, Baek JB, Kim HK (2014) Edge-carboxylated graphene nanoplatelets as oxygen-rich metal-free cathodes for organic dye-sensitized solar cells. Energy Environ Sci 7:10441052. https://doi.org/10.1039/c3ee43732a

Kruusenberg I, Matisen L, Jiang H, Huuppola M, Kontturi K, Tammeveski K (2010) Electrochemical reduction of oxygen on double-walled carbon nanotube modified glassy carbon electrodes in acid and alkaline solutions. Electrochem Commun 12:920923. https://doi.org/10.1016/j.elecom.2010.04.021

Li Q, Wu J, Tang Q, Lan Z, Li P, Lin J, Fan L (2008) Application of microporous polyaniline counter electrode for dye-sensitized solar cells. Electrochem Commun 10:12991302. https://doi.org/10.1016/j.elecom.2008.06.029

Li GR, Wang F, Jiang QW, Gao XP, Shen PW (2010) Carbon nanotubes with titanium nitride as a low-cost counter-electrode material for dye-sensitized solar cells. Angew Chem Int Ed Eng 49:36533656. https://doi.org/10.1002/anie.201000659

Li G, Chen X, Gao G (2014) Bi2S3 microspheres grown on graphene sheets as low-cost counter-electrode materials for dye-sensitized solar cells. Nanoscale 6:32833288. https://doi.org/10.1039/c3nr06093d

Lyu H, Gao B, He F, Ding C, Tang J, Crittenden JC (2017) Ball-milled carbon nanomaterials for energy and environmental applications. ACS Sustain Chem Eng 5:95689585. https://doi.org/10.1021/acssuschemeng.7b02170

Lyu H, Gao B, He F, Zimmerman AR, Ding C, Huang H, Tang J (2018a) Effects of ball milling on the physicochemical and sorptive properties of biochar: experimental observations and governing mechanisms. Environ Pollut 233:5463. https://doi.org/10.1016/j.envpol.2017.10.037

Lyu H, Gao B, He F, Zimmerman AR, Ding C, Tang J, Crittenden JC (2018b) Experimental and modeling investigations of ball-milled biochar for the removal of aqueous methylene blue. Chem Eng J 335:110119. https://doi.org/10.1016/j.cej.2017.10.130

Magnacca G, Guerretta F, Vizintin A, Benzi P, Valsania MC, Nistic R (2018) Preparation, characterization and environmental/electrochemical energy storage testing of low-cost biochar from natural chitin obtained via pyrolysis at mild conditions. Appl Surf Sci 427:883893. https://doi.org/10.1016/j.apsusc.2017.07.277

Musameh M, Wang J, Merkoci A, Lin Y (2002) Low-potential stable NADH detection at carbon-nanotube-modified glassy carbon electrodes. Electrochem Commun 4:743746. https://doi.org/10.1016/s1388-2481(02)00451-4

Nagaraju G, Lim JH, Cha SM, Yu JS (2017) Three-dimensional activated porous carbon with meso/macropore structures derived from fallen pine cone flowers: a low-cost counter electrode material in dye-sensitized solar cells. J Alloys Compd 693:12971304. https://doi.org/10.1016/j.jallcom.2016.10.015

Shao L-L, Chen M, Yuan Z-Y (2014) Hierarchical porous carbons as a metal-free electrocatalyst of triiodide reduction for dye-sensitized solar cells. J Power Sources 272:10911099. https://doi.org/10.1016/j.jpowsour.2014.09.028

Vikrant K, Kim K-H, Ok YS, Tsang DCW, Tsang YF, Giri BS, Singh RS (2018) Engineered/designer biochar for the removal of phosphate in water and wastewater. Sci Total Environ 616:12421260. https://doi.org/10.1016/j.scitotenv.2017.10.193

Wan S, Wu J, Zhou S, Wang R, Gao B, He F (2018) Enhanced lead and cadmium removal using biochar-supported hydrated manganese oxide (HMO) nanoparticles: behavior and mechanism. Sci Total Environ 616-617:12981306

Wang S, Gao B, Li Y, Creamer AE, He F (2017) Adsorptive removal of arsenate from aqueous solutions by biochar supported zero-valent iron nanocomposite: batch and continuous flow tests. J Hazard Mater 322:172181. https://doi.org/10.1016/j.jhazmat.2016.01.052

Xu SJ (2016) Photovoltaic properties of 9 natural leaves derived biochars as counter electrodes for dye-sensitized solar cells. J Chem Ind Eng (China) 67:48514857. https://doi.org/10.11949/j.issn.0438-1157.20160753

Yoon K, Cho D-W, Tsang DCW, Bolan N, Rinklebe J, Song H (2017) Fabrication of engineered biochar from paper mill sludge and its application into removal of arsenic and cadmium in acidic water. Bioresour Technol 246:6975. https://doi.org/10.1016/j.biortech.2017.07.020

You S, Ok YS, Chen SS, Tsang DCW, Kwon EE, Lee J, Wang CH (2017) A critical review on sustainable biochar system through gasification: energy and environmental applications. Bioresour Technol 246:242253. https://doi.org/10.1016/j.biortech.2017.06.177

Yuan Y, Yuan T, Wang D, Tang J, Zhou S (2013) Sewage sludge biochar as an efficient catalyst for oxygen reduction reaction in an microbial fuel cell. Bioresour Technol 144:115120. https://doi.org/10.1016/j.biortech.2013.06.075

Zhang X, Yang Y, Guo S, Hu F, Liu L (2015) Mesoporous Ni0.85Se nanospheres grown in situ on graphene with high performance in dye-sensitized solar cells. ACS Appl Mater Interfaces 7:84578464. https://doi.org/10.1021/acsami.5b00464

Zhao X, Liu W, Cai Z, Han B, Qian T, Zhao D (2016) An overview of preparation and applications of stabilized zero-valent iron nanoparticles for soil and groundwater remediation. Water Res 100:245266. https://doi.org/10.1016/j.watres.2016.05.019

This work was partially supported by the Key Laboratory of Original Agro-Environmental Pollution Prevention and Control, Ministry of Agriculture/Tianjin Key Laboratory of Agro-environment and Safe-product [18nybcdhj-5 and 18nybcdhj-1], the National Natural Science Foundation of China (41807363), Guangxi Natural Science Foundation (Nos.AD17195058), the Key Research and Development Project of the Ministry of Science and Technology (2018YFB0605101), and the Key Project Natural Science Foundation of Tianjin (18JCZDJC39800).

Key Laboratory of Original Agro-Environmental Pollution Prevention and Control, Ministry of Agriculture/Tianjin Key Laboratory of Agro-environment and Safe-product, School of Energy and Environmental Engineering, Hebei University of Technology, Tianjin, 300401, China

Key Laboratory of Pollution Processes and Environmental Criteria (Ministry of Education), Tianjin Engineering Center of Environmental Diagnosis and Contamination Remediation, College of Environmental Science and Engineering, Nankai University, Tianjin, 300350, China

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