Dev. Chem. Eng. Mineral Process., 5(3/4),pp.155-182. 1997. Modelling and Dynamic Simulation of an Industrial Rotary Dryer C. Duchesne', J. Thibault' * and C. Bazin2 D e p a m n t of Chemical Engineering and 2Departmentof Mining and Metallurgy, Lava1 University, Sainte-Foy (Quebec) CANADA GlK 7P4 1 Rotary dryers are widely used in chemical and mineral processes due to theirflexibility and simplicity of operatwn. Nevertheless, the comprehension of the basic phenomena taking place in this type of dryer is still of grea interest. f n order to bener understand the phenomenological behavwur of the process and to assess product moisture control strategies, a complete dynamic simulator of a mineral concentrate dryer hus been developed ;Ilte simulator is calibrated with industrial &a and consists of a furnace model, a solidr traaportaion model and a gas model, including the heat and m s transferphenomena. The dryer simulator was shown to adequately represent the dynamic behavwur of an industrial rotary dryer. Introduction In mineral processing plants, grinding and flotation circuits require water to perform the separation of the valuable minerals from the ore. The product, a wet concentrate sold for its metal content, has to be de-watered to satisfy smelting specifications. Although it tends to be replaced with more efficient equipment, the combination of a vacuum filter followed by a rotary dryer is still widely used. Rotary dryers are simple to operate and are versatile units, but the solids transportation as well as the heat and mass transfer in the rotary drum are very complex phenomena that are influenced by a large number of parmeters. The modelling of these phenomena is still attracting considerable attention in order to operate rotary dryers more efficiently. Therefore, an experimental investigation has been initiated to develop a complete dynamic * Authorfor correqondence. C. Duchesne, J. Thibault and C. Bazin simulator for an industrial dryer, in order to optimise the operating conditions as well as to assess various moisture control strategies. The solids transportation as well as the heat and mass transfer in rotary dryers and kilns has spurred the interest of many researchers for decades. The first significant contributions were experimentally oriented (Friedman and Marshall, 1949; Saeman and Mitchell, 1954; McCormick, 19621, in order to understand the effect of the operating conditions on two important variables, which are the average residence time of solids and the volumetric heat transfer coefficient. Those works contributed to the development of several correlations, which are still being used today. More theoretical approaches were used by Myklestad (1963) and Kamke and Wilson (1986). These authors used differential equations to describe the drymg process and to develop static simulators which were validated on both small and large-scale rotary dryers. Later, more elaborate one-dimensional (Barr et al., 1989; Perron et al., 1992) and three-dimensional (Bui et al., 1995) rotary kiln simulators have been developed for troubleshooting and to solve design problems. Recently, some dynamic rotary dryer simulators have become available (Douglas et al., 1992a and 1992b; Wang et al., 1993). which consist of a set of differential equations combined with correlations for the calculation of the residence time and the volumetric heat transfer coefficient. Even though these models take into account the dynamic behaviour of a rotary dryer, they were only validated on static industrial data. The present paper reports a different approach for the dynamic simulation of an industrial rotary dryer. A solids transportation model has been successfully developed in a previous work (Duchesne et al., 1996). This solids transportation model consists of a series of perfectly mixed interacting tanks with dead volumes, instead of using an empirical correlation. This model was calibrated on an industrial dryer residence time distribution (RTD) and takes into account the effect of lifters, air drag and rheology of the material. These three factors are difficult to model theoretically and are rarely considered in the literature. The heat and m a s transfer phenomena were modelled with differential equations and the heat and mass transfer coefficients were adjusted in order to represent the industrial data as closely as possible. 156 Modelling and dynamic simulation of an industrial rotan d n e r This paper includes the industrial rotary dryer and the data acquisition procedure, followed by the mathematical description of the complete dryer simulator. Finally, the calibration of the model is considered and both dynamic and static performance evaluation, and a sensitivity analysis of the simulator is performed. Equipment and Experimental Procedure The industrial data used for the model calibration and validation was provided by Brunswick Mining and Smelting, New Brunswick, Canada. The company processes ore from a massive sulphide ore-body in New Brunswick and operates a lead sulphide smelting plant at the same location. Four marketable concentrates are produced: zinc, lead, copper and bulk lead-zinc sulphides. These concentrates are thickened and de-watered in vacuum filters, followed by rotary dryers, than shipped to various smelting plants. (a) Concentrate Dryer Brunswick Mining and Smelting operates five concentrate rotary dryers. Figure 1 presents a schematic view of one of the concentrate dryers. Horizontal Enertec or Bigelow-Liptak furnaces (2 m outside diameter and 3.50 m long) are connected to the rotary cylinder. The furnace walls consist of 11 cm of refractory bricks and 11 cm of insulation blocks. The outer shell is made of 1.3 cm of carbon steel. Air is blown into the furnace by two fans and is heated by burning No. 6 fuel oil (Bunker C). The prlmary air and oil are preheated to 50 and 120°C respectively in order to facilitate the combustion. The slope and the speed of the Ruggles-Coles rotary cylinder (2 m outside diameter and 15 m long) are 5 degrees and 5.8 rpm respectively. Straight 15 cm lifters are bolted to the internal surface of the 1.3 cm thick carbon steel rotary drum to facilitate the solids axial displacement and to enhance the air-solid contact. The dryer feed is produced by thickening and filtering the concentrate. The filter discharge is conveyed to the dryer by a belt and a screw conveyor at a rate varying from 15 to 40 metric tonnes per hour. The feed moisture content varies from 15 to 17% by weight and the discharge product moisture content ranges from 7 to 8% by weight. A fan takes the exhaust damped gas to a wet scrubber where dust is recovered and recycled back to the process. 157 C. Duchesne, . I . Thibault and C. Bazin (b) Industria1 Data The data used for the model calibration was obtained during two days of operation, kom zinc dryers #1 and #3, and each day will be referred to as Day 1 and Day 2 data files. Data consists of time series of measurements listed in Table 1. The primary and secondary air flow rates to the furnace and the dryer feed moisture were measured manually for Day 1 only. All the other measurements were provided each minute by sensors installed on the process as indicated in Figure 1. The manual measurements were obtained during the 8 hour test, and some appropriate averages were used as inputs for calibrating and validating the simulator. A ratio controller adjusts the ratio of primary air flow rate to oil flow rate in a proportion that is approximatelyequal to the stoichiometric ratio. The addition of the secondary air provides complete combustion within the furnace, ensures a better control of the gas temperature entering the rotary dryer, and leads to the desired gas flow in the rotary cylinder. Model Development The rotary dryer simulator consists of three sections : a furnace model, a solids transportation model and a gas model. The furnace is an essential part of the complete simulator because two of its inputs are the main manipulated variables of the system. The solids transportation model, which has been developed and discussed in a previous paper (Duchesne et al.,1996) is of utmost importance, because the residence time of the solids within the dryer has a major impact on the drying performance, and the prediction of many important process variables is sensitive to its value (Douglas et al., 1993). A pseudo-stationary gas model was used to simulate the drying medium. Finally, the heat and mass transfer between the solid phase and the gas phase were considered. (a) Furnace Model The furnace provides the hot gas that flows through the rotary dryer and is primarily to evaporate the water. It has three inputs : the primary and secondary air flows and the fuel oil mass flow rate. Preheated air and oil are mixed and burned. As a result of the fuel oil analysis, the stoichiomeuic reaction is described by the following equation based on the assumption of complete combustion : Modelling and dynamic simulation of an industrial rotary dryer Figure 1. Schematic representation of the industrial rotary dryer. Table 1. List of industrial measurements and respective sensors. Process variable I Units I Sensor Dryer feed rate T / h (meuic) Scale Dryer discharge rate T / h (meuic) Scale Exhaust gas temperature OF Thermocouple Feed moisture content % Scaledry-scale Furnace temperature OF Thermocouple Ib I h Corriolis type mass flow meter ACFM Hot wire anemometer % Conductivitymeter ACFM Hot wire anemometer Oil mass feed rate Primary air flow rate Product moisture content Secondary air flow rate 159 C. Duchesne, J. Thibault and C. Bazin The overall heat and mass balances used to calculate the mass flow rate and the temperature of the gas mixture within the furnace are given by the following two equations: mpa +msa +mf = m, ...(2) The residence time of the hot gas in the furnace is small. Indeed, the gas spends approximately0.7 second in the furnace enclosure. It was therefore assumed that the dynamics of the gas in the furnace is negligible. However, to consider the dynamics of the furnace per se, a total heat exchange term has been included in Equation (3) to account for the heat transferred between the gas and the furnace enclosure by convection and radiation. For simplicity, in the calculation of the dynamics of the furnace, the temperature of the refractory walls was lumped and its variation with time was represented by the following first order differential equation : ...(4) Equations (3) and (4) represents the variation of the gas temperature when changes in the air and oil flow rates occur. For instance, if the oil flow rate is suddenly cut to zero, the gas temperature would drop progressively instead of assuming the temperature of the inlet air. The total heat transfer coefficient was obtained by calibrating with industrial data and is given by the following simple correlation : ...(5 ) where 8 and y were determined to be equal to 1 . 4 8 ~ 1 0and ~ 0.613 respectively. This heat transfer coefficient accounts for all heat transfer mechanisms within the furnace. The heat lost to the environment (Qm) has been calculated by considering individual thennal resistances and the overall temperature gradient according to the following equation : 160 Modelling and dynamic simulation of an industrial rotary dryer The total heat coefficient (convection and radiation) at the external surface of the furnace was calculated using available correlations (Incropera and DeWitt, 1990). The thermal conductance (kbJof the insulation layer, between the refractory wall and the steel shell, is not known. In addtion, in a Bigelow-Liptak furnace, the secondary air flows through the narrow cavity between the refractory brick and the insulation block walls. Therefore, its value was estimated so that the outside temperature of the steel shell, under normal operating conditions, would equal the experimentally measured value. The inside refractory wall temperature (Tw) is calculated r fom Equation (4). It is important to point out that this temperature was lumped in Equation (4) in order to account for the dynamics of fhe furnace, and was used as the refractory surface temperature for the calculation of the heat loss. The conduction resistance due to the refractory block insulation is by far the dominant resistance. The enthalpy of a material flow (H)is defined as the weighted sum of the specific enthalpies of all flow components : ...(7) The exit furnace gas is composed of five species : carbon dioxide (COz), nitrogen (N9, oxygen ( 0 2 ) , sulphur dioxide (SOz) and water vapour (H20). The variation of the heat capacity with temperature for each species was obtained from Smith and Van Ness (1987). To calculate the furnace exit gas temperature, the energy balance given in Equation (3) is solved using Newton's method. (b) Solids Transportation Model The solids transportation model is described in detail in Duchesne et al. (1996). The modified Cholette-Cloutier arrangement, as shown in Figure 2, represents the 161 C. Duchesne, J. Thibault and C. Bazin residence time distribution (RTD) of the solids within the industrial dryer. This model is composed of a series of interactive perfect mixers. The volume occupied by the solids in a cell is divided into two zones : an active zone which contributes to the axial transportation of solids along the dryer and a dead zone where the solids are not exchanged with neighbouring cells but only with the corresponding active portion of the cell, as depicted in Figure 3. In passing through the rotary dryer, some portion of the solids is delayed a various number of times by flowing into the dead zones which leads to the extended tail of the residence time distribution curve. The model is described with four parameters that were calibrated using the results of a tracer test (Duchesne et al., 1996). These parameters are : the number of perfect mixers N ; a conductance coefficient k ; the fraction a of the total volume of solids that is occupied by the active zone ; and a parameter b which corresponds to the steady-state ratio of solids exchange rate between active and dead zones to cell mass feed rate. The values of these parameters are summarised in Table 2. The total wet solids and water mass balances for the i" mixer are : ...(8) ...( 1 1 ) where the flow between two adjacent cells is taking place only through the active zones and is proportional to the difference between the total masses of wet solids in the active zones of the two cells, as expressed by the following equation : 162 Modelling and dynamic simulation of an industrial rotary dryer 1 i-1 i N i+l (i-2) k .....x... B Figure 2. Schematic representation of the modified Cholette-Cloutier model. Parameter N k fs-') a P Values 36 0.794 0.75 1 0.013 4 - - - - Figure 3. Mass exchangesfor cell i of the solids transportation model. Under steady-state conditions, b, and bd are both equal to b. However, these parameters are allowed to vary temporarily following a change in feed flow rate in order to obtain the assumed constant volume ratio between the active zone and the total volume of solids (Duchesne et al., 1996). The rate of water evaporation (R), is proportional to the water vapour partial-pressure difference between the solid boundary layer and the gas phase : I63 C. Duchesne, J. Thibault and C.Bazin It was assumed that the partial pressure of water vapour at the gas-solids interface is equal to the saturation partial pressure evaluated using the solids temperature of the active zone, and it was computed from Wagner’s expression (Reid et al., 1987). The latent heat of vaporisation of water was computed using a polynomial regression analysis on steam table data (Smith and Van Ness, 1987), in the range 0 to 100°C. The volumetric mass transfer coefficient is K& and it was assumed constant along the length of the dryer. The variation in enthalpy of the active zone depends on the inlet and outlet flow rates of the cell, the rate of exchange between active and dead zones, the total heat exchange between gas, solids and refractories, and the quantity of water that is evaporated, as given by the following equation : where Ua is the volumetric heat transfer coefficient which was assumed constant along the length of the dryer. A few correlations were proposed to estimate the volumetric heat transfer coefficient, whereas none were found for the volumetric mass transfer coefficient. In the steady-state regime, most volumetric heat transfer coefficients were reported on the basis of the rotary dryer volume (Friedman and Marshall, 1949; McCormick, 1962; Myklestad, 1963). However, in this investigation, the volumetric heat and mass transfer coefficients were defined on the basis of the local volume of solids in the dryer. This implies that even though the volumetric heat (Ua) and mass (Q) transfer coefficientsare kept constant, the rates of heat and mass transfer are modulated by the axial variation of the solids volume within the dryer. Since the volume of material normally decreases along the dryer (Lebaset al., 1994; Duchesne et al., 1996), it is clear that the transfer rates should decrease because of the smaller transfer area. For the dead zone, the enthalpy variation is only due to the exchange of wet solids with the active zone as expressed by the following equation : 164 Modelling and dynamic simulation of an industrial rotary dryer It is assumed that the dead zone does not exchange heat and mass with the gas phase ;I&i and H,,,i are the enthalpies of the wet solids evaluated as a function of the moisture content and the temperature of the active and dead zones of mixer i. These quantities are defined by Equation (7), with two species : the concentrate (mainly zinc sulphide) and water. Gas Model The residence time of the gas phase in the rotary dryer is significantly smaller than the residence time of the solid phase (Douglas et al., 1993). The average residence time is 4 seconds for the gas compared to 20 minutes for the solids in the rotary cylinder. The gas phase was therefore solved as a plug-flow system, that is as a pseudo-steady-state system. The overall mass and water balances on the gas phase are given by : (c) mg.1.-tit - gJ-1 . +Rw mg.i M g . i = mg,i-l Mg,i-l + R w ...( 16) ...(17) where the gas consists of a dry portion, that remains constant within the rotary dryer, and the evaporated water from the solids. It was assumed that solids entrainment by the gas is negligible. As it flows through the dryer, the hot gas transfers heat to the solids and picks up the evaporated water from the solids. Heat is also lost to the environment. The overall heat balance can rherefore be expressed as follows : and I65 C. Duchesne, J. Thibault and C. Bazin QSL ...(19) where R is an adjustable thermal resistance used to calibrate the thermal losses through the dryer shell. Since the heat and mass transfer are evaluated at the mean gas properties (between the inlet and outlet of a ell), few iterations are needed to converge to the mean temperature within an accuracy of 0.05"C.The gas specific enthalpy in cell i (I&d is also defined by Equation (7) with the five species mentioned previously. The mean heat transfer coefficient from the gas to the walls (hb is calculated from Gnielinski's correlation using a hydraulic diameter (Incropera and DeWitt, 1990). The total heat transfer coefficient (convection and radiation) was evaluated from the same correlation used for the furnace external walls. The gas temperature is computed by Newton's method to satisfy the gas heat balance (Equation 18). Both the mean global volumetricheat and mass transfer coefficients (Ua and K@) were determined in order to fit as closely as possible the experimental data. They are defined per unit volume of solids and assumed constant over the dryer length. The volumetric heat transfer coefficient accounts for all heat transfer effects including convection and radiation within the rotary dryer. Results and Discussion (a) Furnace Model In this section, the furnace model is evaluated as to its ability to predict flow rate, composition and temperature of the gas stream leaving the furnace and entering the rotary dryer. In practice, the secondary air flow rate is not measured. Only the fuel flow rate and the gas temperature, remrded via a thermocouple in the furnace enclosure, are normally available on-line. The primary air flow rate is set to a fixed ratio based on the oil mass flow rate according to stoichiometry. With these three variables and with the furnace model described by Eguations (1) to (7), it is possible to estimate the secondary air flow rate that would completely define the exit gas from the furnace. Occasionally. the secondary air flow rate was measured manually, periodically during Day 1 of the sampling campaign. In practice, the secondary air 166 Modelling and dynamic simulation of an industrial rotary dryer flow rate is manipulated via a damper but its value is not measured on-line. Figures 4 to 6 show the variation of the fuel flow rate, the measured temperature and the estimated secondary air flow rate respectively. Figure 6 also shows the measured secondary air flow rate and the stoichiometric primary air flow rate. The estimated secondary air flow rate is significantly lower than the measured flow rate, due to measurement errors of the gas temperature which are discussed in the next section. 0 LL I= 0.00 Q, 0 10000 20000 30000 Time ( s ) Figure 4. Variation of the furnace oil mass~owrate for Day 1. 200 - -- lndusuial D a Q Predicted Data Corrected Oata 800 400 10000 20000 30000 Time ( s ) Figure 5. Comparison between the measured and predicted furnace temperatures. 0 The h a m thermocouplehas no radiation shield to prevent its direct exposure to the flame thermal radiation. Therefore, it was expected that the recorded temperature would be higher than the actual gas temperature because the radiative heat exchange between the flame and the thermocouple cannot be compensated by the convective heat transfer between the thermocouple and the gas stream (Saeman and Mitchell, I67 C. Duchesne, J. Thibault and C. Bazin 1954; Kreith, 1965). To evaluate the measurement error due to radiation, it was necessary to measure the secondary air flow rate in order to calculate the actual gas temperature. With this flow rate (from Figure 6) and the furnace model, a more realistic gas temperature was calculated (compared to Figure 5). in order to satisfy the energy balance for the furnace. These results clearly show the significant bias in the gas temperature which is due to the radiative heat exchange. Because the secondary air flow rate is not currently measured, in order to use the model on the actual process it is necessary to find a way to correct the recorded temperature fiom the measured process variable. Hence estimate as accurately as possible the temperature of the gas entering the rotary dryer. - h rn -4 Y Y Q) c m U : - U .< O L - I Measured SA I-x r-4 <L -'brA'lb. J I -. Calculated SA c - ,. .... . - .. - - - - . . ... . . . - - ._......... .. PA * * . I 0 I I 1 I 10000 1 1 1 20000 Time ( s ) 1 1 1 1 30000 Figure 6. Plot of the ptimaty and secondary airflow rates and total gas flow rate. The heat transfer phenomena taking place within the furnace chamber are very complex. The presence of a luminous flame, the combustion products (H20 and C02) participate in the absorption and re-emission of the thermal radiation, and the evaluation of the various emissivities certainly contribute to the complexity of the problem. To rigorously solve this problem would require considerable effort. In this case, it is desired to simply estimate as accurately as possible the exit gas temperature by correcting the thermocouple reading and adding some dynamics to the f m c e model. The thermocouple receives a net radiative heat flux from its surroundings and looses heat to the gas stream by convection. A higher gas velocity reduces the temperature error due to radiation. Saeman and Mitchell (1954) argued that with a 168 Modelling and dynamic simulation of an industrial rotary dryer very high gas velocity, the temperature measurement error would be negligible. It is therefore not surprising that in Figure 5 the difference between the measured and estimated gas temperatures is smaller at higher gas flow rates. To estimate this temperature difference, an energy balance was performed on the thermocouple where only the convective heat transfer, and the net radiative heat flux between the flame and the thermocouple are considered : The adiabatic flame temperature was dculated and the flame emissivity was calculated as the combined emissivities of H20 and C02,evaluated at the flame temperature using the method proposed by Hottel and Sarofm (1967). The was correlated as a function of the Reynolds convective heat transfer coefficient (h~) number. Since Equation (20) does not give an exhaustive representation of all the heat transfer modes encountered in the furnace, and the values of the flame temperature and emissivity are not well known, the convective heat transfer coefficient accounts for all these modelling errors. The gas temperature estimated with Equation (20) is presented in Figure 5, and on this basis, a correction of 154°C was applied. In practice the secondary air flow rate is not measured, so that Equation (20) would be used to converge simultaneouslyon the total gas flow rate and the gas temperature with the knowledge of the erroneous thermocouple reading. The corrected gas temperature, with Equation (ZO), compares very well to the furnace temperature calculated using the measured secondary air mass flow rate, which is not influenced by radiation. Most of the observed discrepancies may be attributed to the changing behaviour of the flame temperature with the oil flow rate. It is interesting to note that by performing the temperature correction to compensate for thermal radiation, the energy imbalance on the dryer was reduced fkom 92.7% to 3.4% which indirectly validates the procedure used for correcting the gas temperature. The wall of the furnace provides a certain degree of thermal inertia that smooths the gas temperature variation, due to a change in the oil flow rate or the secondary air flow rate. These dynamics were simulated with Equation (3),and the wall temperature is presented in Figure 5. Under steady-state conditions, the model assumes that the wall temperature would be equal to the gas temperature although 169 C. Duchesne. J. Thibault and C. Bazin the wall temperature would actually be higher because of the thermal radiation from the flame. (b) Dynamic Behaviour of Dryer The calibration of the simulator was performed by adjusting the volumetric heat and mass transfer coefficients (Ua and Kgi) and the heat losses, to produce a dynamic response as close as possible to the industrial measurements when subjected to the same inputs. The two main measured variables with which the two transfer coefficients could be calibrated are the solids moisture content and the gas temperature at the exit of the rotary dryer. Since it was felt that the discharge moisture content is the most crucial for the operation of the dryer, it was decided to calibrate the simulator with this variable, and to use the gas temperature as an indirect validation of the simulator calibration. The values of the two transfer coefficients were therefore determined in order to minimise the sum of squares of the difference between the predicted and measured discharge moisture contents. The values of Ua and Q were found to be respectively 10 kW/m3K and 1 kg/s m3kPa. The thermal losses through the dryer shell were calibrated by adjusting the resistance R in Equation (22) in order to obtain the estimated average temperature of 70°C on the outside surface of the rotary dryer shell. Figure 7 shows the variation of the concentrate feed flow rate and the measured and predicted concentrate discharge flow rates for Day 1. A good agreement exists between the predicted and measured values of the &scharge flow rates. These results clearly show that the solids transportation model adequately represents the solids flow behaviour within the rotary dryer. The discharge flow rate shows a Iarge fluctuation that cannot be simulated with the measured concentrate feed flow rate. This stochastic behaviour is characteristic of the transportation of wet solids where sudden local accumulation of material within the rotary dryer can oc(w,followed by the rapid tumbling of material towards the exit of the dryer. Overall, the discharge flow rate is well predicted. The largest deviation is observed when the concentrate feed flow rate was drastically reduced (at around 100 min) which may suggest that for flow rates significantly different from calibrations, the solids transportation model is not able to adequately represent the solids flow behaviour within the rotary dryer. However, it is important to point out that the inlet concentrate feed rate in that 170 Modelling and dynamic simulation of an industrial rotan dryer time span should have led to higher hscharge flow rate. Therefore, this large deviation may also be due in part to the stochastic behaviour discussed above since the lowest and highest discharge flow rates are observed during this t h e period. -u) 10 , r 1 Industrial Data Predicted Data 0 0 20000 10000 30000 Time ( s ) Figure 7. Plot of the solids massjlow rates - Day 1 The solids transportation model was further evaluated with the data set of Day 2 which has a much larger concentrate feed flow rate. Results of this simulation are presented in Figure 8. The concentrate feed flow rate, in addition to being much higher than for Day 1, shows significant fluctuations. Alternatively, the measured discharge feed flow rate is much smoother than that for the previous case. Considering that the solids transportation model was calibrated on a different rotary dryer and that the concentrate feed flow rate varies significantly with time, the solids transportation model provides satisfactory representation of the dynamics of the solids behaviour within the rotary dryer. However, there are some difficulties when 171 C. Duchesne, J. Thibaulr and C. Bazin the concentrate flow rate is significantly higher or lower than the flow rate for which the solids transportation model was calibrated. In order to make the model more robust over a wider range of flow rates, b e parameters of the solids transportation model would have to be expressed as a function of the solids mass flow rate. This could be achieved by performing additional tracer tests at different concentrate feed flow rates. 4 y" 15 g Y al 10 B 0 i Y 5 TI al 2 0 UJ Industrial Data 15 Predicted Data 3 0, x v Q) z U 10 s 0 - u Q) P Q r. .-: : o 5 --I I 1 1 ) 1 I l I J 1 1 1 1 U 0 10000 20000 30000 Time ( s ) Figure 8. Plot of the solids massfIow rates - Day 2. The prediction of the discharge concentrate moisture content for Day 1 is presented in Figure 9, with the measured values. The prediction is good for both the dynamics and the level of the solids moisture content. This was expected since the two transfer coefficients were fitted with these data points. The relatively high variance of the data is mainly due to the sensor used to measure the moisture m t e n t of the solids. The largest deviation is again observed when the feed flow rate was 172 Modelling and dynamic simulation of an industrial rotaq d n e r drasticallyreduced. This time, it is difficult to evaluate the real value of the deviation because the sensor was outside its calibration range. Indeed, during eleven minutes, the recorded value was equal to the minimum value that the sensor could record. h Induslnal Data Prbduted Data a 0 10000 20000 Time ( s ) Figure 9. The measured and predicted product moisture contents. 30000 When the feed flow rate dropped, the predicted moisture content almost decreased to zero. However, in practice, it should level off at a higher moisture Content. Indeed, when the concentrate reaches the critical moisture content, the temperature of the solids rises and the drying rate falls off rapidly. In addition, the drylng rate should approach zero at some equilibrium moisture content which is the minimum value that can be achieved with a particular solid with the drying conditions prevailing within the dryer (Foust et al., 1960). The falling drying rate is normally obtained for moisture contents in the range of 2-3% (Mujumdar, 1987). To include this phenomenon in the simulator, it would be necessary to perform some dryrng experiments to determine the drylng curve at lower moisture contents. However, this does not invalidate the simulator for normal operation. The variation of the moisture Content is well represented by the model. A better prediction could certainly be achieved if the inlet moisture content was measured continuously instead of being set to a constant value of 16.5%. The comparison between the predicted and measured outlet gas temperatures is presented in Figure 10 for Day 1. This comparison can be used as an indirect validation of the complete simulator since it was calibrated strictly on the concentrate discharge moisture content and not on the outlet gas temperature. Apart 173 C. Duchesne, J. Thibault and C. Bazin ffom the drastic change that occurred around 5000 sec, the simulator is able to predict the gas outlet temperature with an average error of approximately6.5"C.This is very good considering that the gas temperature is reduced from about 600°C to 8VC, that is the average temperature deviation is slightly less than 2% of the total temperature drop of the gas. This deviation can have many sources, such as : (i) the small residual energy imbalance of the industrial data ; (ii) measurement errors ; (iii) larger heat losses ; (iv) the assumption of constant volumetric heat and mass transfer coefficients ; and (v) air leakage from outside. - 150 20000 Time ( s ) Figure 10. Measured and predicted discharge gas temperatures. 0 10000 Prduted Data 30000 The plant operating personnel indicated that air infiltration at the discharge end of the rotary dryer is probably the main cause of this difference. As shown in Figure 1, three fans are used to manage the gas flow : the primary and secondary air fans for which the flow can be controlled, and the scrubber fan for which the flow is not controlled. Depending on the gas flow rate through the dryer, the gage pressure at the end of the dryer m a y be negative. Air infiltration through the discharge seals of the rotary dryer can then occur, resulting in a decrease of the measured outlet gas temperature. For a lower gas flow rate, the gage pressure would be lower and the infiltratian of air would be more important. This is corroborated by the results of Figure 10. A rough calculation, using the geometry of the seals and the solids output port, has shown that air infiltration could realistically account for a major component of the observed gas temperature discrepancies. I 74 Modelling and dynamic simulation of an industrial rotan dryer (c) Volumetric Heat and Mass Transfer Coefficients The volumetric heat and mass transfer coefficients are both the product of two distinct parameters : either the global heat transfer (U) or mass transfer(K& coefficient and the solids interfacial area (a). Since the interfacial area is unknown, only the product of the two parameters can be estimated. In this investigation, constant values of Ua and K,+ along the dryer were assumed. However, it is most liely that the two transfer coefficients and the solids interfacial area vary with the operating conditions prevailing within the rotary dryer. A direct relation between the volumetric interfacial area (a) and the volume of solids should not be expected. Basic principles suggest that the interfacial area should increase with a decrease in the moisture content because of the lower compaction and the higher bed void associated with a lower moisture content. The area should also increase with a decrease of the solids volume because a higher proportion of solids would be directly exposed to the hot gas. The heat and mass transfer coefficients (U and KJ should realistically increase with an increase in the gas flow rate. However, increasing the gas flow rate leads to a reduction of the residence time of the solids within the rotary dryer, resulting in a reduction of the contact time between the gas and the solids, thereby opposing the beneficial effect of higher transfer coefficients. In view of the above considerations, it would be difficult to derive a universal correlation for predicting the volumetric heat and mass transfer coefficients. The assumption made in this investigation of constant volumetric heat and mass transfer coefficients appears reasonable and practical. (d) Static Application of the Simulator Results presented in the previous sections demonstrated the ability of the simulator to realistically represent the dynamic behaviour of an industrial rotary dryer. To assess the simulator under steady-state conditions, a simulation was performed with a concentrate feed flow rate of 6.87 kg/s, oil flow rate of 0.057 kg/s, primary and secondary air flow rates of 0.754 and 2.79 kg/s. The resulting temperature of the gas entering the rotary dryer was 686°C. The temperature and the moisture content profiles of the gas and the solids as a function of the position along the dryer are presented in Figure 11. I 75 - C. Duchesne, J. Thibault and C. Bazin a I- v c c a C t 10 6 2 3 c zj I- .- 100 v) 0s 0 0 5 10 15 Axial Position ( m ) Figure 11. Axial profiles of the temperatures and moisture content. Both gas and solids temperature profiles are consistent with those reported by Friedman and Marshall (1949) which used dry material. Most of the heat is transferred to the solids within the first 5 m of the dryer. According to Saeman and Mitchell (1954), it is usually observed that 90% of the heat transferred in a murrent flow rotary dryer OCCUTS within the first 10 % of the dryer length. Near equilibrium conditions for heat and mass transfer are obtained at about 5 m from the entrance and the outlet solids temperature is close to the gas wet-bulb temperature. Therefore, the gas and solids moisture contents are almost constant in the last 10 m of the dryer. The exhaust gas is only at one half of its water saturation, showing that the heat and mass gradients between the gas and the solids, are limiting the heat and mass transfer. Therefore, it would be interesting to reduce the sensitivity of the operating conditions in order to reach the desired moisture content near to the discharge end, rather than at one-third of the dryer length. This type of study may results in considerable fuel economy. (e) Sensitivity Analysis The sensitivity of important simulated process variables, with respect to the input variables and parametersof the model, are of utmost importance in the interpretation of predictions and in the design of future sampling campaigns. Furthermore, a sensitivity analysis can assist in the selection of manipulated variables for controlling and optimising the process. The relative sensitivity of each variable and parameter was evaluated using the following sensitivity index : 176 Modelling and dynamic simulation of an industrial rotan)d q e r ..(21) The sensitivity index was determined by changing one parameter at the time by 10%.and then by performing a complete simulation using operating conditions of Day 1. The values Y and Yrefcorrespond to the output variable after the change and its reference value respectively, whereas aef is the nominal value of the modified parameter and w is the new value used to generate a new value of the output variable being considered. Three process variables and fourteen parameters were changed in turn and their effect was observed on four dryer outputs, namely: the moisture content and temperature of both the outlet gas and solids streams. The results are presented in Table 3 and the sensitivity of the various parameters and variables can be easily assessed. The sensitivity analysis of the three input process variables (secondary air, fuel and concentrate feed mass flow rates) shows that the secondary air flow rate is ten times less sensitive than the other two input variables. The logical control strategy calls for the manipulation of the oil mass flow rate since the concentrate feed flow rate cannot be manipulated. In addition, the sensitivity analysis shows that in the region of the nominal operating conditions, the secondary air flow rate should be as small as possible. Therefore, the sensitivity analysis of the input variables suggests adjusting the fuel flow rate in order to suppIy sufficient energy necessary to evaporate the water, and to set the secondary air flow rate to give an entrance gas temperature as high as possible while satisfymg mechanical and thermal constraints. This strategy is consistent with the strategy used by the operators at Brunswick Mining and Smelting. The sensitivity of the volumetric heat transfer coefficient is much more significant than the sensitivityof the volumetric mass transfer coefficient. This result confirms that, under normal operating conditions, heat transfer controls the drying process.Tbe three parameters that have the most influence on the concentrate discharge moisture content are: the heat of combustion, the latent heat of vaporisation, and the inlet solids moisture content. The latent heat of vaporisation is known, and therefore for correct simulation, an accurate value of the fuel oil heat of 177 C.Duchesne, J. Thibaulr and C. Bazin Table 3. Sensitiviv index (S)of the four main process outputs. T40 M, MW -0.176 0.010 -0.043 -0.570 0.205 - 1.250 0.497 0.695 -0.147 1.578 -0.480 0.480 -0.139 2.698 -0.342 -0.455 0.011 0.000 -0.053 0.000 -0.217 0.000 0.036 0.000 0.002 -0.006 0.002 0.003 0.254 -1.317 0.507 0.882 0.003 -0.010 0.004 0.006 -0.017 0.083 -0.027 -0.060 0.020 -0.106 0.036 0.072 0.005 -0.023 0.009 0.014 0.025 -0.033 0.033 0.082 0.0oO 0.001 0.000 -0.002 -0.214 0.970 -0.306 -0.692 0.011 -0.001 0.012 0.034 -0.281 -0.123 -0.240 0.119 combustion should be available. As expected the feed moisture content greatly influences the discharge moisture content. For a more accurate calibration of the simulator, and if a feedforward control strategy is to be implemented, it would be necessary to measure the inlet moisture content on-line. In the present investigation, it was assumed constant and equal to 16.5%. The relatively high sensitivity index suggests that the assumption of constant inlet solids moisture content may be 178 Modelling and dynamic simulation of an indusrrial rorary dryer responsible for some of the observed deviations between experimental and simulation results. Conclusions A dynamic simulator for an industrial rotary dryer was presented. The mathematical models included the production of combustion gas in a furnace, solids transportation within the romy dryer, gas flow and heat and mass transfer phenomena. The simulator was calibrated with a limited number of experimental data gathered over a period of 8 hours. The solids transportation model uses interactive mixers in series with, 25% of the volume occupied by the dead zone. This model is able to adequately represent the flow of solids through the rotary dryer under both steady and transient conditions, provided that the mass flow does not deviate too far fiom normal operating conditions. The volumetric heat and mass transfer coefficients were then determined in order to minimise the prediction error on the discharge solids moisture content. The outlet gas temperature was predicted satisfactorily and served as an indirect validation of the complete simulator since the two transfer coefficients were not adjusted to minimise the deviations of this variable. A sensitivity analysis has provided the relative influence of the input process variables and model parameters on the temperature and moisture content of the gas and solids at the discharge of the dryer. Results suggest an appropriate discharge solids moisture control strategy for the moisture content of the discharging solids. The results also indicated the importance of measuring the feed solids moisture content for better calibration of the simulator, and for possible implementation of a feedforward control strategy. Despite limited measurement of some process variables, it is believed that the described simulator provides a good representation of the various phenomena taking place within an industrial rotary dryer. This investigation developed and calibrated a rotary dryer simulator based on dynamic industrial data. In future sampling campaigns, it would be desirable to determine the influence of concentrate and gas flow rates on the residence time distribution curve, because the solids transportation model has a major impact on all phenomena taking place within an industrial rotary dryer. I 79 C.Duchesne. J. Thibault and C.Bazin Acknowledgements The authors would like to thank a consortium of eight mining companies under the umbrella of CAMIRO, NSERC, Centre de Recherches Minerales du MER and CANMET for their support of the KBAC research program. Special thanks goes to Brunswick Mining and Smelting for their co-operation during the sampling campaign. Nomenclature Solids interfacial area Cross sectional area Specific heat capacity Fuel specific combustion heat Mean total heat transfer coefficient Heat transfer coefficient to environment Total enthalpy of a stream Flow conductanceof wet solids Volumetric mass transfer coefficient Dryer length Furnace length Mass of material Moisture content (wet basis) Mass flow rate Number of chemical species in a stream Number of dryer slices Water partial pressure in the flue gas Saturation water vapor-pressure Heat losses Reynold's number Thermal resistance used in Equation (27) Radius Water evaporation rate Sensitivity index Time Temperature Ambient air temperature Volumetric heat transfer coefficient Wet solids volume Componentmass fraction 180 kg/s Pa Pa kW Modelling and dynamic simulation of an industrial rotary dryer Greek letters a Volume of active zone to the total solids volume aT Absorptivity of the thermocouple Mass flow rate exchange ratio between active and dead zones P Corrected mass flow rate fraction entering a dead zone Pa Corrected mass flow rate fraction going out of a dead zone Pd eF Emissivity of the flame eT Emissivity of the thermocouple in Equation ( 5 ) Coefficient g lw Water latent heat of vaporisation (kJ1kg) (kW I m2 K) Coefficient in Quation (5) 9 Air density kg I m3) rair (W I m2 K") S Stefan-Boltunannconstant W Variable in Equation (21) Variable in Equation (21) Y Subscripts a bics C cs d eCSW efw f F FL fw f3 i icsw ids in 1 0 Pa rbi rw S sa SL T V W Active zone Interface block insulation - carbon steel Concentrate Carbon steel Dead zone External carbon steel wall External furnace wall Fuel Flame Furnace losses Furnace wall Gas Indicates properties associated with the itb cell Internal carbon steel wall Inside dryer shell Input Indicates a species or acts like a counter in a summation Outlet Primaryair Refractory block insulation Refractory wall solids Secondary air Shell losses Thermocouple Vapour Water 181 C. 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