Determination of power requirements for solid core pulp screen rotors

213-217 4703 04-04-16 16.34 Sida 213 Determination of power requirements for solid core pulp screen rotors James A. Olson, University of British Columbia, Canada, Serge Turcotte and Robert W. Gooding, Advanced Fiber Technologies (AFT) Inc., Canada KEYWORDS: Pulp, Screening, Rotor, Experimental, Power SUMMARY: Pilot and mill scale pulp screening trials were conducted to determine the effect of rotor design and operation on the power required by a solid core rotor. The pilot plant experiments were conducted over a wide range of rotational speeds, feed flow rates, rotor element shapes, and reject ratios, using 3% softwood kraft pulp. The mill scale trials were conducted in two mills using two different size screens. The resulting power required was analyzed and given in terms of the nondimensional power coefficient, capacity coefficient and Reynolds number. From the experimental trials, we were able to show that the power required by the rotor increases approximately as tip-speed cubed and rotor diameter squared. Furthermore, it was shown that the dimensionless power required increased linearly with dimensionless screen capacity for all of the pilot scale rotors tested. The rate at which power increased with capacity was related to the shape of the hydrodynamic elements attached to the solid core. In addition, the dimensionless power - capacity curve was shown to determine the power requirements for a given rotor design and to provide an important basis for comparing the power requirements of different rotors. ADDRESSES OF THE AUTHORS: James A. Olson (olson@mech.ubc.ca): The Pulp and Paper Centre, the Department of Mechanical Engineering, The University of British Columbia, 2324 Main Mall, Vancouver BC Canada V6T 1Z4. Serge Turcotte(serge.turcotte@aft-global.com): Advanced Fiber Technologies (AFT) Inc., 72 Queen Street, Lennoxville, Quebec, Canada, J1M 2C3. Robert W. Gooding (robert.gooding@aft-global.com): Advanced Fiber Technologies (AFT) Inc., 5890 Monkland Avenue Suite 400, Montreal, Quebec, Canada, H4A 1G2. Pressure screens improve pulp quality and value by removing contaminants that degrade the appearance, optical properties and strength properties of paper. Modern pressure screens are also capable of improving pulp quality by fractionating the fibres for selective processing or for use in specialty paper grades. Screening is a critical process in high value, pulp and paper production and its importance has increased further with the development of advanced products and the increased use of recycled furnishes. Modern screens improve quality by passing the pulp fibres through small aperture screen cylinders. In general, the smaller the aperture the larger the product improvement gained. The minimum aperture dimension is determined by the effectiveness of the screen rotor, thus, the rotor is an important component in the pulp quality improvements available in screening. Pressure screens are increasingly required to operate at higher consistencies. High consistency screening reduces the cost of capital equipment by reducing the required installed screen capacity and by reducing the requirements for downstream thickening equipment. The cost savings by moving to high consistency screening are especially realized in mills where low-consistency operations, such as, cleaning are not used. However, at high consistencies the power required by the screen rotor can become a significant cost factor. Rotors impart energy to the pulp to do three things: (1) Disrupt the fibre network at the screen cylinder surface so that the pulp moves as individual fibres through the apertures, (2) accelerate the fluid to a high tangential velocity to improve the probability screening effect and (3) provide a negative pressure pulse that backflushes the apertures and clears them of any fibre accumulations. The effect of rotor design and operation on the efficiency, thickening and fractionation efficiency is complicated. In general, as more energy is applied, the maximum capacity of the screen increases, and the efficiency and reject pulp consistency decreases. The detailed effects of rotor design and operation on screen performance have been extensively reviewed in several recent studies, including Wakelin and Corson (1995), Julien Saint Amand (1999) and Gooding and Olson (2001). In general, it is known that less aerodynamic rotor elements require more power, have higher capacity and lower efficiency. The effect of rotor speed and gap on the magnitude and shape of pressure pulses induced by the rotor was investigated by Pinon, Gooding and Olson (2002). Despite the large number of studies examining the impact of rotor design and operation on screen performance, there are few studies relevant to the power requirements of the screen rotor. However, there have been several studies that examine the power required to fluidize pulp in a rotary device which is a similar application and these studies provide some insight into the power requirements of pressure screens. The power required to fluidize a volume of fluid was experimentally measured using a concentric rotary device by Gullichsen and Harkonen (1981), Bennington and Kerekes (1996) and Bennington, Kerkees and Gracce (1992). In these studies, they measured the critical power dissipation per unit volume for a range of rotor and housing diameters to achieve pulp fluidization. These studies provide an estimate of the minimum power required to fluidize a pulp as a function of the pulp consistency and can provide estimates of the minimum power required to achieve fluidization in a pressure screen. In addition to the minimum power requirements for fluidization, Bennington and Kerekes measured the power required as a function of angular velocity for fluids with a wide range of viscosities and for pulps with a wide range of consistencies. Following the classic approach, 213

213-217 4703 04-04-16 16.34 Sida 214 typically used to describe the power requirements of pump and other rotary equipment, Bennington and Kerekes non-dimensionalized the power and angular velocity using the rotor diameter, D, fluid density, ρ and angular velocity, ω. The resulting dimensionless power and Reynolds number was then defined as (Note: an alternative form of the dimensionless parameters is also given in terms of the peripheral speed of the rotor or tip speed, Vt, which is more commonly used in the application to pulp screens.): In their study, they showed that for small Reynolds number (<10 3 ) the power coefficient decreased roughly linearly with Reynolds number and that for larger Reynolds number (>10 3 ) the power coefficient was approximately independent of Reynolds number. From [Eq 1], we see that a constant power coefficient yields the commonly used industrial rule of thumb that states that the power per square meter of screening area is proportional to tip velocity cubed, assuming geometrical similarity. In the application to high consistency screening, where the suspension has an initial yield stress, we need to assume that the suspension is sufficiently fluidized such that it behaves as a Newtonian fluid with an apparent viscosity, µ a. In addition to the above dimensionless variables, pumps and other rotary equipment typically described the fluid flow rate through the device, Q f, using the dimensionless capacity coefficient (White 1999), defined as In the application to pulp screens, the diameter is the rotor diameter and the angular velocity is the angular velocity of the screen (Fig 1). Fig 1. (a) A generic solid core rotor depicting the variables used in this analysis and (b) a photo of an industrial solid core rotor. In this study, we experimentally measure the power required to operate a pressure screen with a solid core rotor. The experiments are conducted over a wide range of rotor speeds, feed flow rates, rotor sizes, rotor element shapes, and reject ratios. The results are presented in terms of the dimensionless parameters defined above. Both pilot and mill screening studies are presented. Experimental A series of pilot screening trials were carried out using the pilot scale, Centrisorter M200 located at Vancouver laboratory of the Pulp and Paper Research Institute of Canada. The screen is outfitted with pressure and flow sensors on the feed, accept and reject ports. The feed pulp was an Eastern Canadian, fully-bleached, softwood, kraft pulp at 3% consistency and the screen plate was a MacroFlow(tm) MF1240 slotted screen cylinder with a 0.25 mm wide slot, a 1.2mm high contour and 4.0mm wide wires. The rotor is attached to a variable frequency drive which allows a large range of rotor velocities to be tested. The power drawn by the motor is measured using a power meter attached to the rotor drive. Three solid core rotors (Fig 1) were used in this study, each with a different element shape. The rotors element shapes are shown in Fig 2 and are referred to as a Gladiator HC (GHC) element, an LR element and an experimental element named the XR. All have significantly different element shapes and represents the common elements used in industry. Fig 2. Schematic drawing of the three solid core element shapes used in this study. (a) an LR element, (b) a Gladiator HC (GHC) element and (c) an Experimental (XR) element. For each element shape, power consumption was measured as a function of (1) increasing rotational velocity, typically ranging from 1000 to 1500 RPM, (2) increasing feed flow rate at a constant angular velocity and volumetric reject ratio and (3) increasing volumetric reject ratio at a constant accept flow rate and angular velocity. Mill trials were carried out on two geometrically similar Centrisorter screens, located at two separate mill locations, to determine the effect of rotor diameter on power consumption. These trials, including the pilot trials, provided three diameters of rotors, equal to 0.38, 0.5 and 0.8 m for the M200, M400 and M800 screens. In these tests, the same variable frequency drive (VFD) was installed at each mill to vary the rotational speed of the screens. The VFD also provided a means of measuring the rotor power consumed. The measured and reported power includes the losses from the bearings and the motor inefficiencies. The M400 screen was located at an OCC mill and was used as a primary coarse screen. The feed flow was approximately 3500 lpm at 3.5% consistency and the reject flow rate was controlled to be approximately equal to 700 lpm. The screen cylinder was a 2 mm diameter, 0.9mm contoured (A4 Profile ), drilled hole. The accept flow through the screen was controlled to be constant approximately constant. 214

213-217 4703 04-04-16 16.34 Sida 215 The M800 screen was also located at an OCC mill and was used as a secondary screen in the fine screening system. The feed pulp was a 1.0% consistency and the feed flow rate was approximately 7000 lpm. The volumetric reject flow was held constant at 860 lpm with a consistency of approximately 1.6%. The screen cylinder was continuous, 0.2mm wide slotted basket with a 0.9 mm contour and 4.2 mm slot pitch (i.e., MacroFlow MF 0940). Results and Discussion Effect of rotor speed and element shape The effect of element shape and rotor speed was determined by varying the rotational speed of the M200 pilot pressure screen and measuring the power consumed for all three element shapes. The volumetric reject ratio and feed flow rate were held approximately constant for all tests. Fig 3 shows the increase in power with increased Fig 3. Power consumption as a function of rotor tip-speed at a constant R v and Q f for all three element shapes tested. tip speed for the three element shapes tested. A least squares fit to a power function was performed. The resulting exponent for the XR and LR elements was calculated to be approximately 2.83 while the GHC exponent was calculated to be 3.1. The correlation coefficient for the least square fits was greater than 0.99 for all three rotors. On average, the power required by the rotor increases approximately as the third power of tip-speed. Fig 4 shows the same data plotted in non-dimensional form, i.e., power coefficient and Reynolds number. From Fig 4, it is evident that the power coefficient, C p, is approximately independent of Reynolds number. This indicates that the flow in the screening zone is fully turbulent and that the energy losses are independent of the viscosity of the fluid, i.e., the power coefficient is independent of the Reynolds number. The independence of Reynolds number corresponds with the experiments of Bennington and Kerekes (1996), despite the large differences in rotor geometry. We note that the Reynolds number in the pilot screening experiments is greater than 10 7, assuming that the apparent viscosity of pulp is approximately equal to water. Even if the apparent viscosity of the feed pulp suspension was 1000 times that of water (which it can be for 10% consistency pulp) the Reynolds number will still be greater than 10 4 for all experiments would still be greater than the critical Reynolds number for fully turbulent behaviour, according Fig 4. Dimensionless power coefficient as a function of Reynolds number. to the experiments of Bennington and Kerekes. Fig 4, also clearly demonstrates the difference in power required for the three rotors elements used in this study. The XR shape was significantly higher than both the LR and the GHC rotors. The GHC had the lowest power consumption of the rotors tested. It is evident that the element shape significantly impacts the power required. This implies that to minimize power consumption, it is important to have a low-drag element shape like the GHC element. Effect of feed flow rate and reject ratio The effect of feed flow rate on required power was determined by varying the feed flow at a constant volumetric reject ratio and tip speed for all three element shapes in the M200 pilot pressure screen and measuring the consumed power. Fig 5 shows the dimensionless power coefficient plotted as a function of the dimensionless capacity coefficient. From this figure, it is evident that the power consumed by the rotor increases linearly with the feed flow rate. Using a linear regression and extrapolating power to zero flow conditions (i.e., C q =0) enables a comparison of the three rotors at zero flow rate. The power coefficient, Cp, at zero feed flow is approximately equal to 1.5, 1.5 and 1.8 for the GHC, LR and XR rotor elements, respectively. This value is in close agreement with the power coefficient measured by Bennington and Kerekes in their fluidization experiments where there was no flow of pulp suspension through the rotary device. Furthermore, this indicates that the power required to overcome viscous dissipation, at zero flow into the screen, may be approximately independent of the geometry of the rotor elements. Fig 5. Power coefficient as a function of capacity coefficient the rotor elements tested. 215

213-217 4703 04-04-16 16.34 Sida 216 From Fig 5 it is evident that the different rotor designs result in a different slope between C p and C q. The difference may be due to the differing ability of the rotor elements accelerate the incoming feed pulp, i.e., the higher slope corresponds to a rotor element shape that induces a higher tangential velocity in the fluid. diameter rotors require more power than smaller diameter rotors to operate. A least squares analysis indicates that the increase in power with increasing tip speed is approximately given by a cubic exponent. The two industrial screens (M400, M800) were found to have a slightly lower exponent, equal to 2.7, while the pilot screen (M200) exponent was 3.1. Fig 6. Power coefficient as a function of capacity coefficient for varying volumetric ratio. Fig 6 shows the power versus capacity coefficient for a range of volumetric reject ratio, R v =0.1 to 0.4. We see that the increase in power consumption is approximately independent of the volumetric reject ratio, despite the change in reject pulp consistency from 3.0% to over 4% at the lowest R v. The lack of dependence on reject consistency is probably due to the fact that the consistency only increases to 4% in a small annulus at the bottom of the cylinder. Effect of rotor diameter (mill trials) To determine the effect of rotor diameter, experimental trials were carried out on a Centrisorter M400 (0.5m diameter) and an M800 (Diameter = 0.8m diameter). The M400 was located at a O.C.C. mill and the M800 was at an OCC mill. Both screens were outfitted with a Gladiator HC rotor, AFT MacroFlow slotted for M800 and drlled 2mm holes for M400 screen cylinder and a variable frequency drive. The power was measured for a range of increasing rotor speeds. The feed flow rate was held approximately constant for each test but varied significantly for each screen. However, it is still possible to compare the power requirements because the power required by the Gladiator HC screen was shown to be relatively independent of feed flow rate (as shown in Fig 5). Fig 7 shows the measured power versus tip speed for both the mill screens and the M200 pilot screen. Fig 7. Power consumption as a function of RPM for constant R v and Q f. for all three Centrisorter M-Series screens. From Fig 7, it is evident that, at constant tip speed, larger Fig 8. Dimensionless power coefficient as a function of Reynolds number (dimensionless angular velocity) for three different Centrisorter sizes and installations. Fig 8 shows the non-dimensional power coefficient versus Reynolds number for the three Centrisorter screens. In general, the power coefficient for each screen is independent of Reynolds number and is approximately the same for all three screens, despite the differences in mill installation, furnish and consistency. However, we do note that the M800 screen has a lower power requirement than the other screens. The difference is most likely due to the significantly lower consistency of the feed pulp in the M800 trial. A reduction in consistency from 3.5% to nearly 1% resulted in about a 15% reduction in dimensionless power coefficient. In any event, the differences are relatively small, considering the large range of rotor diameters and rotational velocities used in this study. Conclusions In this study, the power required by a solid core rotor was experimentally determined for a wide range of rotor speeds, feed flow rates, rotor sizes, rotor element shapes, and reject ratios. The experiments were carried out under realistic (mill and pilot) conditions with acceptable screening efficiencies. From the experimental trials we were able to show that: 1. The power required to operate a pressure screen rotor was shown to be approximately proportional to the tip speed cubed. 2. The experimentally measured power coefficient was shown to be approximately constant for a wide range of rotational speeds and rotor diameters. This agreeswith the commonly used industrial rule of thumb which states that the power per meter squared of screening area increases with the tip velocity cubed. 3. The power coefficient was shown to increase linearly with the capacity coefficient and the slope was strongly dependent on the geometry of the rotor elements. The elements with blunt leading edges had a higher slope than the aerodynamic element 216

213-217 4703 04-04-16 16.34 Sida 217 shape because blunt leading edge elements accelerate more of the feed flow to a high velocity. The Gladiator HC rotor was found to be nearly independent of feed flow rate, whereas, the less aerodynamic element shapes (e.g., with blunt leading edges) showed a strong dependence on the feed flow rate. In summary, this study demonstrates that the power coefficient - capacity coefficient curve determines the power requirements for a screen rotor and provides an important basis for comparing the power requirements of different rotor designs. The study also demonstrates how to predict the power consumption required by a screen rotor and allows higher precision engineering of new screening systems and system upgrades. Acknowledgments The authors would like to acknowledge the technical assistance of Norm Roberts and Phil Allan of the Pulp and Paper Research Institute of Canada (Vancouver laboratory). Literature Julien Saint Amand, F., and Perrin, B., Fundamentals of screening: Effect of rotor design and fibre properties, 1999 TAPPI Pulping Conference Gooding, R., and Olson, J.A., Parameters for Assessing Fibre Fractionation and their Application to Screen Rotor Effects International Mechanical Pulping Conference, Helsinki, June 2001 Pinon, V., Gooding, R.W., and Olson, J.A., Measurements of pressure pulses for a solid core rotor Tappi Pulping/Engineering Conference, San Diego, 2002 Bennington, C.P.J. and Kerekes R.J., Power requirements for pulp suspension fluidization Tappi J. 79(2) 1996. Gullichsen, J., and Harkonen, E., Tappi 64(6):69, 1981 Wakelin, R., and Corson S., TMP Long fibre fractionation with pressure screens Int. Mech Pulping Conf., 1995. Bennington C.P.J., Kerekes R., and Grace J., The Yield Stress of Pulp Suspensions CJChE, 1992 White, F.M., Fluid Mechanics 4th Edition, WCB McGraw-Hill, 1999 Manuscript received September 3, 2003 Accepted February, 2004 217