Biomass screw feeding with tapered and extended sections

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Powder Technology 186 (2008) 56 – 64 www.elsevier.com/locate/powtec

Biomass screw feeding with tapered and extended sections Jianjun Dai, John R. Grace ⁎ Department of Chemical and Biological Engineering, University of British Columbia, 2360 East Mall, Vancouver, British Columbia, Canada V6T 1Z3 Received 8 August 2007; accepted 31 October 2007 Available online 17 November 2007

Abstract Successful feeding is critical to biomass utilization processes such as gasification and combustion, but feeding is difficult due to the adverse effects of heterogeneity, physical characteristics and moisture content of the particles. The objectives of the present study were to define blockage mechanisms and plug seal with the aid of tapered and extended sections. Wood pellets, sawdust, hog fuel and wood shavings were delivered by a screw feeder/lock hopper system. The experimental results showed that tapered and extended sections increased torque requirements significantly while improving the plug seal to the reactor, with performance depending on biomass physical properties and screw feeder characteristics. © 2007 Elsevier B.V. All rights reserved. Keywords: Biomass; Feeding; Screw feeder; Torque; Taper; Extended section

1. Introduction Interest in using biomass feedstocks to produce heat, power, liquid fuels and hydrogen, as well as to reduce greenhouse gas emissions, is increasing worldwide. Biomass processes, including direct combustion, gasification and pyrolysis, have been under development for many years. A critical problem is how to feed biomass into the reactors. The major problems for biomass feeding are blockage and plug seal failure to the reactor [1]. Fuel feeding problems often impede smooth operation in industry. If the reactor operates at high-pressure and/or high temperature, there are additional challenges in establishing reliable feeding [2–5]. Feeding problems often impede smooth operation. Such particle properties as mean size, size distribution, shape, surface roughness (e.g. smooth, rough or sharp edges), density, moisture content, compressibility, and pliability can all affect the success of feeding. In biomass energy processes, several kinds of feeders and their combinations have been reported, in particular hopper or

⁎ Corresponding author. Tel.: +1 604 822 3121; fax: +1 604 822 6003. E-mail address: [email protected] (J.R. Grace). 0032-5910/$ - see front matter © 2007 Elsevier B.V. All rights reserved. doi:10.1016/j.powtec.2007.10.033

lock hopper systems, screw feeders, rotary valves, piston feeders and pneumatic feeders. These types can also be combined, especially for continuous operations. These feeders can handle a variety of solids, but they have limitations with certain types of biomass and/or in feeding to pressurized reactors. Screw or piston operated plug feeders, common in feeding coal, have also been tested with biomass [6–13]. Roberts [14,15], Roberts and Manjunath [16], and Roberts [17] analyzed the volumetric characteristics and mechanics of screw feeders in relation to the bulk solid draw-down characteristics of the feed hopper. Distribution of throughput along screw and uniform draw-down patterns was investigated. Yu and Arnold [18] proposed a theoretical model for torque requirements for single screw feeders. They assumed that the load imposed on a screw feeder by the bulk solids in the hopper is determined by the major consolidation stress. Considering the bulk material boundary in a pocket between adjacent flights, forces are imposed on five surfaces. Compression in the choke section was neglected due to the short length of the choke section. Biomass screw feeders operate in a manner similar to piston feeders, but have a somewhat lower pressurization range (0.5– 1.5 MPa). In a screw feeder, an auger compresses the feedstock into a compact plug, aided by tapering the feed channel, or by gradually reducing the pitch of the screw. The feed plug then

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Fig. 1. Geometries of lower hopper and screw feeder.

forms a barrier, preventing backflow of gases and bed material from the reactor [8,12]. The length of the choke section for biomass feeding is typically much longer than for more conventional materials such as coal, due to the requirements of plug seal formation. In this paper, the effects of the choke section, including tapered and extended sections, on screw feeding are investigated. A methodology is presented to analyze the tapered and extended sections employed to enhance plug sealing to the reactor. Mechanisms of blockage and prediction of torque requirements for biomass screw feeding with tapered and extended sections are delineated. Predictions are then compared with experimental results. Temperature and thermal effects are neglected in the present study. 2. Experimental set-up, methodology and materials fed The biomass feeding systems used in the experiments appear in Figs. 1 and 2. Tapered and extended sections are shown in

Fig. 1. Table 1 gives key dimensions of the hopper-screw feeder. Wood pellets, sawdust, hog fuel and wood shavings served as biomass materials. Polyethylene particles provided a reference material for comparison. The main physical properties of the particulate materials are listed in Table 2. Biomass particles were added to the lower hopper and the surface was then flattened. A variable-speed DC motor (0.56 kW, Baldor CDP 3440) adjusted the rotational speed of the screw. A weighing scale (B.C. Scale Co. Ltd., Model: CARDINAL EF 100, capacity: 50 kg, accuracy: 0.02 kg) connected to a computer recorded the weight of material fed at 2 s time intervals. A torquemeter (S. Himmelstein and Company, Model: MCRT28004T 5-3, capacity: ± 565 N m, accuracy: ≤± 0.1% of full scale) with Model 721 Mechanical Power Instrument was installed between the DC motor gear reducer and lower hopper to measure the screw torque and speed of rotation at a frequency of 36 Hz during feeding. Blockages were detected by changes in torque and rotational

Fig. 2. Schematic of biomass feeding system.

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Table 1 Hopper and screw a dimensions Screw

Trough

Hopper

Clearance, c Flight thickness, γ Average helix angle of screw with vertical Material Inside diameter, Dti Material Transparent test section Type Length Height Angle of hopper wall with horizontal Material

force, a balance of forces acting on an element of length dx results in 1, 6 and 11 mm 6.35 mm 14° 316 SS 102 mm Carbon steel Cast acrylic tube Wedge-shaped 910 mm 910 mm 70° Carbon steel

rx kr2 ¼ ðrx þ drx Þkðr  dx tan at Þ2 " # kr2 kðr  dx tan at Þ2 þsw  tan at tan at h i þrw kr2  kðr  dx tan at Þ2 : Elimination of second order terms yields drx 

a

Screw diameter 100 mm for length of 800 mm, then 90 mm for length of 300 mm, and finally 80 mm for length of 420 mm.

speed. All data were stored in a data acquisition computer for later analysis. Each experiment was performed 2–5 times in order to determine the range of flow rates and torque values for a given material and corresponding experimental conditions. All measurements were carried out at room temperature (20 ± 2 °C) and atmospheric pressure. The fill level in the feed hopper declines during each feeder trial unless the feeder is periodically refilled. The flow rate was determined from the average flow rate in the first 2 min after stable feeding was established. Feed rate and torque readings are time-averaged for analysis. When blockage occurred and flow could not be re-established, the materials were removed from the screw feeder and hopper manually before the next run. Three initial hopper levels were tested: high (0.60 m), medium (0.45 m) and low (0.30 m). 3. Mechanics and torque analysis for hopper-screw feeders There are two main regions for the screw feeder, the hopper feeding section and choke section (or conveying section) as shown in Fig. 1. Further details of the feeder load and stress analysis in the choke section are provided elsewhere [1]. 3.1. Taper casing in choke section Tapered sections of length 0.15 and 0.30 m at the discharge end of the screw feeder were tested in the present study (Fig. 1). The outlet diameter of both tapered sections was 88 mm, so the taper angles for the two tapered sections differ, as shown in Fig. 3. Taking axial resistance forces caused by the tapered casings into account increases the torque requirements. Volumereducing flow channels (i.e. converging tapered sections) block more readily for incompressible materials (such as polyethylene and wood pellets) unless the feeder is powerful enough to break the particles. For compressible bulk materials (e.g. sawdust and ground hog fuel), blockage is difficult to estimate. It depends on the material properties and flow conditions inside the screw casings. The geometry affecting the stress of a material element within a pocket in the tapered section is shown in Fig. 4. Assuming steady state for a moving material element and equilibrium between the driving force and resisting

ð1Þ

2rx tan at 2sw 2rw tan at dx þ dx þ dx ¼ 0: r r r

ð2Þ

Substituting r = (Dt − 2xtanαt) / 2 into Eq. (2), we obtain drx þ

2rx 2sw 2rw dr  dr ¼ 0: dr  r r tan at r

ð3Þ

After substituting σw / σx = λs, τw = μwσw and rearrangement, Eq. (3) becomes   drx 2ks Aw dr : ð4Þ ¼ 2  2ks r rx tan at Boundary conditions for the 0.30 m long tapered section (P = 0.1 m) are: ◦ first pocket: σx = σin at r = Rt ◦ second pocket: σx = σin1 at r = Rt1 ◦ third pocket: σx = σin2 at r = Rt2. Boundary condition for the 0.15 m long tapered section: ◦ first pocket: σx = σin at r = Rt ◦ second half pocket: σx = σin1 at r = Rt1. Here Rt is the inside radius of the casing at x = 0, Rt1 the inside radius of the casing at x = P, and Rt2 the inside radius of the casing at x = 2P (only for 0.30 m long tapered section), as shown in Fig. 4. σin, σin1, σin2 are axial stresses at the starting point of successive gaps in the tapered section. Rt is used in the following sample calculations. Integrating Eq. (4) and substituting r = (Dt − 2x tan αt) / 2 yields   x tan at h rx ¼ rin 1  ð5Þ Rt where h ¼ 2ð1  ks  ks Aw = tan at Þ: The average axial stress within a pocket is   Z 1 P x tan at h rin 1  dx: rxa ¼ P 0 Rt

ð6Þ

ð7Þ

Hence the average normal wall stress in a pocket is rwa ¼

ks P

Z

P 0

  x tan at h rin 1  dx: Rt

ð8Þ

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Table 2 Key properties of materials in the present study Name of specimen Polyethylene particles Wood pellets Ground wood pellets-1 Ground wood pellets-2 Sawdust-1 Sawdust-2 Sawdust-3 Hog fuel-1 Hog fuel-2 Hog fuel-3 Ground hog fuel Wood shavings-1 Wood shavings-2 Wood shavings-3 a b

Mean diameter a (mm)

Size range (mm)

Bulk density (kg/m3)

Particle density (kg/m3)

Moisture (wet basis) b

4 9.8 4.05 0.55 0.45 0.45 0.45 0.72 0.72 0.72 0.18 0.67 0.67 0.67

3.0–5.0 8.0–11.6 3.35–4.75 100% b 3.35, 98.5% N 0.09 100% b 6.73, 96% (0.09 to 2.8) 100% b 6.73, 96% (0.09 to 2.8) 100% b 6.73, 96% (0.09 to 2.8) 100% b 25, 90% (0.09 to 9.5) 100% b 25, 90% (0.09 to 9.5) 100% b 25, 90% (0.09 to 9.5) 100% b 4.75, 98.7% (0.09 to 2.8) 100% b 12.5, 91% (0.09 to 6.73) 100% b 12.5, 91% (0.09 to 6.73) 100% b 12.5, 91% (0.09 to 6.73)

610 630 485 423 210 330 440 200 310 322 150 110 156 188

908 1200 1200 1200 370 550 688 360 490 510 330 300 380 430

dry 8% 8% 8% 14% 40% 60% 11% 40% 60% 14% 10% 40% 60%

Sauter mean particle diameters except for the first two which are volume-equivalent diameters. Measured by drying at 105 °C for 5 h in an oven.

For both the 0.15 and 0.30 m long tapered sections, the compression can be estimated by stress analysis in the choke section [1] and by the volume changes along the screw axis in the tapered sections. For each pocket in a tapered section, the screw pocket volume can be calculated, and then used to estimate the compacted bulk density. Average axial stresses within each pocket in the taper section for various materials can be estimated by an axial stress–bulk density relation, ρy = a(σy + b)c. The stress σin can be obtained from Eq. (7) by iteration until the calculated average axial stress σxa matches the calculated compacted bulk density. The torque requirement generated within each pocket in the tapered section can be estimated from a force analysis [1]. The 0.15 m tapered section covers one-and-ahalf pockets, whereas the 0.30 m tapered section includes three pockets. Total torque requirements for the tapered sections are obtained by summing the torque required for each pocket.

(1) 0.30 m long tapered section The axial resistance acting on an element of bulk solids on the core surface can be expressed by Z Fca ¼ 2kRc rwa Awc sin ðac Þ

P

dx 0

¼ kcd cp rwa Awc sin ðac Þrwa D2o :

ð9Þ

The axial resistance force acting on the material element in a pocket on the trough surface is expressed by Fta ¼ ð 2kRPAwt rwa þ 2kRPrwa tan at Þ

ð10Þ

where R is the inside radius of the taper section on the driving side of the flight, i.e. at the beginning of each pocket (at x = 0, P and 2P for the 0.30 m long tapered section).

Fig. 3. Dimensions of two tapered sections tested in this work.

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Fig. 4. Stresses on material element in tapered section.

The axial force acting on the material within a pocket on the trailing side is Z Ro cos ð/f  ar Þ Ffa ¼ 2kks rxf rdr Rc cos ar cos /f Z Ro ¼ 2kks rxf ð1 þ tan ar tan /f Þrdr: ð11Þ Rc

After substituting tan αr = P / 2πr and tan ϕf = μf, Eq. (11) can be solved numerically or analytically. The total axial force on the material element within a pocket caused by the driving side of a flight should equal the total resisting axial forces acting on the same material element due to the core shaft surface, trailing side of a flight and trough surface. It is again assumed that the total force is uniform on the surface of the driving side, so that Fda þ Fca þ Ffa þ Fta ¼ 0

ð12Þ

Fda : rda ¼  2 k Ro  R2c

ð13Þ

The torque generated by the core shaft surface is Tc ¼ 2kR2c PAwc rwa cos ac :

ð16Þ

The torque generated by the trailing side of the screw flight is obtained by numerical integration of Z Ro Z sin ð/f  ar Þr2 dr 2k Tf ¼ ks rxf dh cos ar cos /f Rc 0 Z Ro ð tan /f  tan ar Þr2 dr: ð17Þ ¼ 2kks rxf Rc

The torque generated by the flight tip is Ttip ¼

grwa PAf Ro : sin ao

ð18Þ

The normal stress on the flight tip surface is assumed to be σwa. The flight tip surface area within a pocket can be estimated by Atip = γ × P / sinαo.

The torque generated in one pocket is Ti ¼ Td þ Tc þ Tf þ Ttip

ð14Þ

where Td is the torque due to driving flight, Tc is the torque due to core shaft surface, Tf is the torque due to trailing flight, Ttip is the torque due to the flight tip, N m. The torque generated by the driving side of the screw flight is Z Ro Td ¼ 2krda r2 tan ðar þ /f Þdr: ð15Þ Rc

After substituting tan αr = P / 2πr and tan ϕf = μf, Eq. (15) can be integrated numerically.

Fig. 5. Comparison of torque predictions and experimental measurements for 0.15 m tapered section and 0.45 m initial hopper level.

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3.2. Extended section in choke section In order to propel bulk materials forward in the extended section, the axial force exerted by the driving side of the screw flight must equal or exceed the resisting force, i.e. kDt Le rewa Aw V

Fig. 6. Comparison of torque predictions and experimental measurements for 0.30 m tapered section and 0.45 m initial hopper level.

The total torque generated from the tapered section is the summation of the torque required by each pocket in the tapered section. The torque predictions for hopper feeding section and straight choke section are described by [1]. (2) 0.15 m long taper section The 0.15 m long taper section includes one-and-a-half pocket (P = 0.1 m). The average axial stress in the second half pocket is   Z 2 P=2 x tan at h rxa ¼ rin 1  dx: ð19Þ P 0 Rt1 Hence the average normal wall stress in a pocket is   Z 2ks P=2 x tan at h rwa ¼ rin 1  dx: P 0 Rt1

ð20Þ

The trailing side of the flight is absent in the second half pocket for the 0.15 m tapered section. The force and torque calculations in this case are based on P/2 instead of P.

kD2t rexa 4

ð21Þ

where σexa is the average axial stress and σewa is the average normal wall stress in the extended section. Substituting σewa = λsσexa yields

Le VLct1 ¼

Dt : 4ks Aw

ð22Þ

There is another critical length defined by Lct2 ¼ Dt tan d:

ð23Þ

Lct1 and Lct2 are critical lengths for extended sections, and δ is the effective internal friction angle. Note that Dt (casing diameter) is employed instead of Do (screw diameter) in these equations. When the length of the extended section exceeds the critical length, bulk materials cannot be easily pushed out of the extended section since the force exerted by the screw flight is not transmitted forward effectively, but instead transmitted to the wall, tending to cause blockage for incompressible materials and significant compaction for compressible materials. The compression and blockage tendency for

Fig. 7. Stress on material element in extended section.

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Table 3 Recommended critical length of extended section (ID: 102 mm) for various materials

4. Comparison of model predictions with experimental results

Name of specimen

For blockage conditions, the predicted torques are those required to reach certain compression levels inside the tapered and extended sections, not those required to break up a blockage or to advance the plug.

Wall friction angle on Internal friction Lct1 carbon steel (°) angle (°) (m)

Wood pellets Ground wood pellets-1 Ground wood pellets-2 Sawdusts-1 Hog fuel-1 Ground hog fuel Wood shavings-1 Polyethylene particles

31.4 30.2 30.4 31.8 31.5 31.8 31.0 21.5

32.0 33.2 38.0 38.0 39.0 45.0 39.0 26.2

0.09 0.11 0.15 0.14 0.15 0.21 0.16 0.14

Lct2 (m) 0.06 0.07 0.08 0.08 0.08 0.10 0.08 0.05

compressible materials in the extended section depends on the fullness of screw pockets and material properties. When Le NLct1 ¼

Dt 4ks Aw

ð24Þ

or Le NLct2 ¼ Dt tan d

ð25Þ

particular attention should be paid to ensure effective compression and prevent unwanted blockage inside the extended section. The average axial stress in the extended section can be calculated based on the measured compacted bulk density in the extended section. Then the average normal stress in the extended section can be estimated by σewa = λsσexa. From a force balance at steady state (i.e. during blockage or moving at a constant speed), the following expression can also be obtained kDt Le rewa Aw ¼

kD2o rda : 4

ð26Þ

Rearrangement of Eq. (26) gives the stress on the driving side of the screw flight rda ¼

4kDt Le rewa Aw : kD2o

ð27Þ

4.1. Taper casing in choke section Torque predictions and experimental measurements are compared in Figs. 5 and 6. For sawdust and ground hog fuel, the predictions agree well with the experimental results, but the experimental results are much higher than the predictions for wood shavings. Blockage occurred for both the 0.15 and 0.30 m long taper sections for the wood shavings. The large deviations are probably due to the wide size distribution and wide range of particle strength for wood shavings. The torque predictions for hog fuel in the 0.15 and 0.30 m long tapered sections with a hopper level of 0.45 m are 276 and 419 N m respectively, suggesting blockage (motor capacity b 100 N m), although this could not be confirmed in the present experiments. For the 0.30 m long tapered section, blockage always occurred, whereas blockage sometimes, but not always, occurred for the 0.15 m long tapered section. The difference is attributed to the complicated flow patterns in the hopper and in the choke sections, leading to varying levels of fullness in the choke section. Increased fullness tended to cause blockage for both the 0.15 and 0.30 m long tapered section. Although the taper angle for the 0.15 m tapered section was larger than for the 0.30 m tapered section, the greater length made blockage more likely. Screw rotational speeds affected the tendency to block in the tapered section. Higher screw speeds (e.g. 30 and 40 rpm) tended to transport more bulk materials into the tapered section per unit time, increasing the likelihood of blockage compared to lower screw speeds. The torque predictions neglect any effect of screw speed. From observations on blockage of hog fuel during experiments, torque N 100 N m is expected to disrupt blockage inside

In Eq. (27), Do is employed instead of Dt.

Fig. 8. Comparison of torque predictions and experimental measurements for 0.15 m extended section and 0.45 m initial hopper level.

Fig. 9. Comparison of torque predictions and experimental measurements for 0.30 m extended sections and 0.45 m initial hopper level.

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the entrance to the choke section) is recommended for the choke section length. 5. Conclusions Different choke section configurations lead to different compression in the choke section, as well as different torque requirements. As a rule, longer choke sections, tapered and extended sections lead to better plug seal at the expense of increased torque requirements for screw feeding. A methodology is presented to analyze the tapered and extended sections which are employed to enhance plug seal to reactors in the biomass industry. Fig. 10. Comparison of torque predictions and experimental measurements for different choke section lengths.

the tapered section and make feeding smoother. It should be noted that predictions for hog fuel and wood shavings are particularly uncertain for the tapered section cases due to the wide particle size distributions, wide ranges of particle strength and complex flow patterns in the hopper and choke sections. 4.2. Extended section in choke section Stresses on a material element in the extended section are shown in Fig. 7. The torque generated by the extended section can be predicted by the above equations. The recommended critical lengths for various materials are listed in Table 3, while the predicted torques with extended sections appear in Figs. 8 and 9. Although the predicted torques for a 0.15 m extended section for sawdust-1 do not match the experimental data very well, approximate estimates can be provided for biomass materials with consideration of torque fluctuations. Torque readings generally fluctuated in the range 10–30 N m during the experimental tests. The plug formed in the extended section becomes even tighter when the axial stress exerted by the screw flight increases. A tighter plug also increases the resistance exerted by the casing wall, depending on the length of the extended section and the material properties. It is difficult to predict the required torque to expel the plug into the receiving vessel for different biomass materials. The critical length of the extended section is important if a screw feeder with extended sections is employed. It then acts like a piston feeder. Several plug formation regions (i.e. casings without screw flights) along the screw in the choke section are expected to provide better performance from a plug seal point of view [19]. The model developed by [1] for the same screw is intended for different choke section lengths (i.e. 0.30, 0.46 and 0.60 m), as shown in Fig. 10. For a shorter length (e.g. 0.30 m), the compression factor for torque predictions is the same as for a longer choke section (e.g. 0.60 m) in the present experiments. Hence the torque requirements are reduced compared to the longer choke section. Longer choke sections are expected to provide better plug sealing and a greater probability of blockage. For biomass feeding, 4–10 times the screw pitch (at

Nomenclature Atip Flight tip surface area within a pocket, m2. Do Screw flight diameter, m. Dt Inside diameter of trough or casing, m. P Pitch of screw, m. F Force, N. Lct1 Critical length of extended section in Eq. (24), m. Lct2 Critical length of extended section in Eq. (25), m. Le Length of extended section, m. r Screw radius, m. Rc Screw core shaft radius, m. Ro Screw flight radius, m. Rt Trough or casing inside radius, m. Rt1, Rt2 Inside radius of casing at x = P and x = 2P in tapered section, m. Ti Torque generated in one pocket, N m. T Torque, N m. x Coordinate in horizontal direction, m. Greek letters αc Screw flight helical angle at core shaft surface, tanαc = P / 2πRc, radians. αo Screw flight helical angle at outside radius, tanαo = P / 2πRo, radians. αr Flight helical angle at screw radius r, radians. αt Half-angle of tapered section, radians. δ Effective internal friction angle, radians. ϕf Wall friction angle of bulk solids on flight surface, radians. γ Screw flight thickness, m. λs Ratio of normal wall stress to axial stress for bulk solids sliding on surface, –. μf Wall friction coefficient between bulk solids and screw flight surface, = tanϕf, –. μw Wall friction coefficient between bulk solids and trough or core shaft surface, –. θ 2(1 − λs − λsμw / tan αt) in Eq. (6). σda Stress acting on material element in a pocket due to driving flight side, Pa. σin, σin1, σin2 Axial stresses at starting point of successive gaps in tapered section, Pa. σw Normal wall stress perpendicular to trough wall and core shaft surface, Pa.

64

σwa σx σxa σxf σewa σexa τw

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Average normal wall stress perpendicular to trough wall and core shaft, Pa. Axial compression stress inside screw feeder, Pa. Average axial stress in a pocket, Pa. Axial stress on trailing side surface in tapered section, Pa. Average normal wall stress in extended section, Pa. Average axial stress in extended section, Pa. Shear stress on trough wall and core shaft surface, Pa.

Subscripts a Axial direction c Core shaft d Driving side of flight f Trailing side of flight o Outside diameter t Trough surface tip Flight tips w Direction normal to wall x Axial direction y Vertical direction

References [1] J. Dai, J.R. Grace, A model for biomass screw feeding, Powder Technol. 186 (2008) 40–55, doi:10.1016/j.powtec.2007.10.032. [2] T.C. Elliott, Standard Handbook of Power Plant Engineering, McGrawHill, New York, 1989. [3] FBT, Inc., Fluidized bed combustion and gasification: a guide for biomass waste generators, Prepared for Southeastern Regional Biomass Energy Program, 1994. [4] M.A. Cuenca, E.J. Anthony, Pressurized Fluidized Bed Combustion, Blackie Academic and Professional, London, 1995.

[5] K.R. Cummer, R.C. Brown, Ancillary equipment for biomass gasification, Biomass and Bioenergy 23 (2002) 113–128. [6] N. Bundalli, D. Morgan, O. Martinez, Bin-feeder design for reliable flow of hog fuel, Pulp Pap. Can. 87 (1987) 145–148. [7] A. Ghaly, Al Taweel, A. Ergudenler, Development and evaluation of a straw feeding system for fluidized bed gasifier, Canadian Bioenergy R&D Seminar, 1989, Technical University of Nova Scotia Halifax, Nova Scotia, 1989, pp. 297–303. [8] C. Wilén, A. Rautalin, Handling and feeding of biomass to pressurized reactors: safety engineering, Bioresour. Technol. 46 (1993) 77–85. [9] D. Nelson, The development of a screw feeder for hogged bark and other wood refuse, Canadian Pulp and Paper Association, 80th Annual Meeting, Technical Section, 1994. [10] S.P. Babu, Thermal gasification of biomass technology developments: end of task report for 1992 to 1994, Biomass and Bioenergy 9 (1995) 271–285. [11] M. Gabra, H. Salman, B. Kjellstrom, Development of a sugar cane residue feeding system for a cyclone gasifier, Biomass and Bioenergy 15 (1998) 143–153. [12] K.R. Cummer, R.C. Brown, Ancillary equipment for biomass gasification, Biomass and Bioenergy 23 (2002) 113–128. [13] X.T. Li, J.R. Grace, C.J. Lim, A.P. Watkinson, H.P. Chen, J.R. Kim, Biomass gasification in a circulating fluidized bed, Biomass and Bioenergy 26 (2004) 171–193. [14] A.W. Roberts, Determining screw geometry for specified hopper drawdown performance, Proc. Bulk 2000 Conference, Institution of Mechanical Engineers, London, 111–116, Oct., 1991. [15] A.W. Roberts, Basic Principles of Bulk Solids Storage, Flow and Handling, Institute for Bulk Materials Handling Research, University of Newcastle, 1992. [16] A.W. Roberts, K.S. Manjunath, Volumetric and torque characteristics of screw feeders, Proc. Powder and Bulk Solids Conf., Rosemont, USA, vol. 189–208, May 1994. [17] A.W. Roberts, Predicting the volumetric and torque characteristics of screw feeders, Bulk Solids Handl. 16 (1996) 233–244. [18] Y. Yu, P.C. Arnold, Theoretical modeling of torque requirements for single screw feeders, Powder Technol. 93 (1997) 151–162. [19] L. Bates, Guide to the Design, Selection and Application of Screw Feeders, Professional Engineering Publishing, London, 2000.