About this Journal Submit a Manuscript Table of Contents
Advances in Mechanical Engineering
Volume 2013 (2013), Article ID 538931, 12 pages
http://dx.doi.org/10.1155/2013/538931
Research Article

Conjugated Heat and Mass Transfer in Convective Drying in Compact Wood Kilns: A System Approach

Department of Mechanical Engineering, State University of Santa Catarina, Campus Universitário Prof. Avelino Marcante, 89219-710 Joinville, SC, Brazil

Received 19 March 2013; Accepted 3 October 2013

Academic Editor: Hakan F. Oztop

Copyright © 2013 M. Vaz Jr. et al. This is an open access article distributed under the Creative Commons Attribution License, which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.

Abstract

Drying is a required operation in timber manufacturing industries. Most drying processes utilise convective kilns, which involve coupling between transient heat and mass (moisture content) transfer in solids and convection in a flowing medium. In the present work, the flow field in the kiln is determined using a system approach based on a general head loss equation, whereas the coupling between transient heat and mass transfer in the timber is accomplished by a solution of a local differential model across the timber stack. The scheme is applied to study the effects of distinct geometric kiln configurations and air flow temperatures upon drying rates and moisture content.

1. Introduction

The technological advancements in the last years have made available to industries several conventional and nonconventional wood drying methods such as radiofrequency heating, press, solar, and dehumidification drying strategies [1]. Notwithstanding, convective drying is the mostly used method in agricultural, food, paper, and timber processing industries [2]. In general applications, convective drying of moist objects (e.g., lumber manufacturing, food processing, etc.) conjugates heat and mass transfer both inside the solid and in the flowing medium. The understanding of the related physics of the problem is crucial for the process control and ensuing product quality.

The industrial importance of drying processes has been translated by the increasing number of works reporting numerical prediction of temperature distribution and moisture content in convective drying problems. In convective timber drying, temperature, moisture content, air velocity, and drying rates are the most important parameters to affect product quality; that is, an incorrect setting of global parameters may cause defects such as colour heterogeneities, warp, and splits amongst others. Therefore, development of simple and effective predictive models to evaluate the process parameters is of paramount importance in lumber manufacturing industries. Riepen and Paarhuis [3] present a numerical study on the optimization of flow velocity in convection kilns. The authors conclude that there is a strong relationship between drying rate and air velocity distribution inside the kiln. Kaya et al. [46] devise 2D numerical models to predict heat and mass transfer during convective drying. The authors solve the flow problem to obtain the convection heat transfer coefficient in the boundary of the moist object considering a laminar flow regime. Analogy between thermal and concentration boundary layers is adopted to estimate the convection mass transfer coefficient on the surface, followed by solution of the transient diffusion problem (heat and moisture inside the object). In a similar direction, Mohan and Talukdar [2] present a 3D computational model to study the convective drying of a moist object. More recently, two works assessing convective drying of wood have been reported in the literature [7, 8]. Younsi et al. [7] utilise a 3D model to compute the turbulent flow inside convective kilns and evolution of the temperature and moisture content. Thibeault et al. [8] combine the Ansys CFX (heat and moisture transfer) and FESh++ (mechanical approximation) commercial codes to predict the thermo-hygro-mechanical behaviour of wood during convective drying.

Most literature works involving convective drying focus only on the local physics of some particular moist object (small scale). To the authors’ best knowledge, there are no works in the literature based upon a system approach where the convective kiln is modelled in the real (full) scale and air flow is computed using a lumped approximation. It is expected that the kiln geometric configuration, as well as lumber stacking, and sticker number, geometry, and placement affect both the head loss and flow aerodynamics and, as a consequence, the flow operating condition (and, ultimately, the drying rate). Based upon the authors’ experience in modelling and simulation of forced convection heat transfer in industrial problems (e.g., [1012]), a full scale, globally iterative coupled simulation (in conjunction with the fan performance curve to obtain the operation point) of air flow and heat and mass transfer is computationally intensive and, within the industrial environment, would require CFD advanced users to select proper model settings. Therefore, the present system approach aims at developing a simpler model and is yet able to obtain good approximations to the physical phenomena. Following this route, it is proposed a mixed numerical model (a global approach to determine flow velocities and a local differential model to compute evolution of the timber temperature and moisture content) to predict the effects of distinct kiln configurations and air flow temperature on the evolution of drying rates, moisture content, and temperatures in a full geometric scale.

2. System Approach: The Flow Operating Condition

The strategy used in this work to estimate air velocity across the timber stack is based upon determination of the flow operating point of the global system—the system approach. This technique requires (i) knowledge of the fan performance curves—provided by the manufacturer and (ii) computation of the system characteristic curve. The typical system characteristic curve correlates the total head loss (or pressure drop) and flow rate in the inlet/exit plenums and timber stack, accounting for friction losses and effects of flow separation. This section describes the procedure to obtain the characteristic curve of the system and the operating condition for the compact dry kiln depicted in Figures 1 and 2.

538931.fig.001
Figure 1: Compact dry kiln geometry: lumber stack and sticker placement.
538931.fig.002
Figure 2: Compact dry kiln geometry: lateral view.

Air flow between two consecutive timber layers is dictated by the pressure difference between the inlet and exit plenums. Ideally, the kiln design should provide uniform pressure in both plenums, making it possible to obtain the same velocity in all flow channels. However, the plenum configuration or inadequate sticker dimensions/placement may cause pressure differences and, consequently, variations in air velocity across the stack height, which may lead to undesirable, nonuniform drying rates. The present formulation follows the concept of Nijdam and Keey [13], who developed a differential head loss equation to compute channel velocities across timber stacks assuming that the kiln configuration is such to prevent boundary layer separation at the top corner. Aiming at a discrete solution for a known number of timber layers and stickers, the proposed approach determines the flow rate based upon an equivalent flow circuit and accounts for channel inlet and exit effects and flow through parallel channels between stickers.

The geometrical information is indicated in Figures 1 and 2. The main parameters are the stack height, , width, , and length, ; number of stickers and corresponding dimensions, and ; number of lumber layers across the stack height, ; and maximum and minimum inlet and exit plenum widths, and . Such configuration causes air to flow between timber layers through parallel, individual channels of cross-section . The stack level is indicated by (upward direction), in which . The geometry of the inlet and exit plenums, depicted in Figure 2, defines a continuous cross-section variation, , so that and .

The pressure drop, , between entrance and exit sections of a generic channel of length can be readily derived from the mass conservation principle and the classical head loss equation [14], so that where subscripts “in” and “out” indicate entrance and exit sections, is a loss coefficient (when separation takes place in contraction/expansion flows), is the kinetic energy correction factor, is the specific mass, is the cross-sectional area, is the volumetric flow rate, is the total hydraulic resistance, and is the friction factor, approximated using Swamee and Jain’s equation [15]: in which is the channel roughness, is the air viscosity, and is the hydraulic diameter.

It is relevant to notice that (1) is nonlinear, which prevents use of a direct analogy with Ohm’s liner relation between voltage (pressure drop) and electrical current (flow rate). However, a similar approach can be used based upon the mass conservation and by assuming equality of pressure at the entrance/exit of the channels at a given stack level.

2.1. Equivalent Global Flow Circuit

Figures 1 and 2 show that air flows through the inlet plenum, splits into parallel channels at every discrete level of the timber stack, and exits through the exit plenum. The configuration of the equivalent flow circuit is shown in Figure 3, which highlights the flow through parallel channels for stack levels and (Figure 3(a)) and flow through the timber stack and inlet/exit plenum stretches (Figure 3(b)). In view of the present kiln geometry and flow circuit presented in Figure 3, application of (1) requires computation of the hydraulic resistances for the inlet/exit plenum stretches, and (between stack levels and ), and parallel channels, (between stickers for stack level ), as in which the subscripts and represent plenum and individual channel measures and is the stack level across the timber stack height; the superscripts and indicate inlet and exit plenum stretches and denotes an individual flow channel at a given stack level . The hydraulic diameter for a flow channel, , and the corresponding mean values for the inlet/exit plenum stretches, , are where , , , and are indicated in Figures 1 and 2.

fig3
Figure 3: Equivalent flow circuit: (a) channels (between two timber layers) and (b) inlet/exit plenum stretches and channels (across the stack height).
2.2. Individual Channel Approximation

Figure 1 shows that insertion of stickers between two timber layers creates parallel channels with equal inlet and exit pressures, and , respectively, at every stack level . Furthermore, the hydraulic resistance for each individual channel at a given level is constant, , (i.e., equal flow rate, , at each parallel channel). Therefore, solving (1) for the total flow rate, , in the stack level , one obtains and the hydraulic resistance, , for level as

2.3. Inlet and Exit Plenum Approximation

The inlet and exit plenums are split into stretches of equal height , as indicated in Figure 2, for which hydraulic resistances and are computed using (3). Therefore, in addition to and , the global circuit configuration includes also the hydraulic resistance for stack level , , as illustrated in Figure 3(b). The pressure drop in the channels and plenum stretches for two consecutive stack levels, and , are, respectively, in which is the flow rate in the plenum stretch. By combining (7), it is possible to determine a recurrence equation for and as a function of the flow rate for the parallel flow channels and plenum stretches at stack level as

The total flow rate, , and pressure drop, , in the system are determined by acknowledging that the flow rate in level (under the first timber layer) is the same as in the first inlet/exit plenum stretches, . The subsequent values across the stack height , , are determined recursively using (7) and (8). Therefore, the system characteristic curve can be readily determined by assuming different values of within a range , from which the corresponding total flow rate, , and pressure drop, , are determined at the end of the recursive computation.

2.4. Fan Performance Curve and Operating Point

Conceptually, the operating point consists in the balancing flow condition between the mechanical energy provided by fans and the losses imposed by the flow topology in the dry kiln, represented by the point at which the fan performance curve at a given rotational speed intersects the system characteristic curve. Mathematically, the flow operating condition can be determined by the solution of the nonlinear expression , where and are the system and fan characteristic equations obtained by application of nonlinear (polynomial) regression to the system and fan characteristic curves as in which (,) and (,) are the maximum order of the regression polynomials and corresponding coefficients for the system and fan, respectively. A second-order polynomial, , for the system would suffice for a coefficient of determination .

The fan characteristic curves are provided by the manufacturer and should indicate the respective rotational speeds; otherwise the so-called fan laws [16] can be used to estimate the performance curves within a limited range of speeds. It is worthy to mention that the fan performance curve may present a nonmonotonic behaviour with respect to the flow rate, which might require a piecewise approximation to improve accuracy. Furthermore, additional care should also be exercised to avoid operation near the fan instability region, that is, where the system characteristic curve intersects the fan curve in more than one point or both curves present similar slopes.

At the operating point, the flow rate, , is computed at the entrance of the inlet plenum using the recurrence equation (8). Therefore, in addition to , the corresponding flow rates, , in each individual channel at stack level are also available, making it possible to estimate the individual channel velocities .

3. Heat and Mass Transfer

The transient heat and mass diffusion processes in a timber layer are modelled by a one-dimensional formulation. The global and local approaches are illustrated in Figures 1, 2 and 4. Figures 1 and 2, show the timber stack and lateral view of the kiln in a full scale, whereas the local configuration is presented in Figure 4, which portraits a detailed view of the local geometry with its main dimensions and variables.

538931.fig.004
Figure 4: Local geometry: flow channels, where , , and are the channel height, length, and depth (the latter is not represented in the picture: see Figure 1), respectively, and timber layers with thickness .

Assuming that (timber layers are thin), the mathematical model comprising the energy and chemical element (moisture) conservation principles reduces, respectively, to where is the direction normal to the gas flow, is the timber temperature, is the thermal diffusivity of the wood, represents the moisture content in the solid (kg of moisture/kg of dry wood), and is the diffusion coefficient of moisture in the timber. The diffusion coefficient, , is temperature dependent, being given by the Arrhenius law in which and are experimental constants. The present work assumes  m2/s and  K, corresponding to the drying wood process in a wide range of temperatures [17].

Computation of the global drying curve requires solution of (10)-(11) for each individual timber layer owing to possible differences in the flow velocities across the timber stack. In addition, the boundary conditions at the solid/gas interface and the initial condition for transient heat and mass transfer problems must also be specified. The energy balance at both interfaces of a timber layer of thickness leads to where is the thermal conductivity of the solid (timber layer), is the air flow temperature, and are temperatures at the solid/gas interfaces, and is the convective heat transfer coefficient. The latter is determined by assuming fully developed flow of the gas in the channels so that [18] for laminar and turbulent flow regimes, respectively. In (13), is the hydraulic diameter, is the thermal conductivity of the gas, and and are the Reynolds and Prandtl numbers, respectively. Therefore, the individual mean channel velocity at the operating point, , (detailed in the preceding section) and definition of the air temperature make it possible to obtain the heat transfer coefficient, (computed individually for each flow channel, ).

Application of the mass balance for the chemical element (moisture) at the air/timber interfaces leads to where is the moisture content of the drying air, and are the moisture content in the timber/air interfaces, and is the convective mass transfer coefficient. The coefficient is determined by using the analogy between thermal and concentration boundary layers, [18], that is, where is the Lewis number and is the diffusion coefficient of the chemical component (moisture) in B (air). It is important to state that the procedure adopted in the present work for modelling conjugated heat and mass diffusion has also been adopted by other authors [2, 46] within similar framework.

The heat and mass conservation equations (10) are fully coupled through the mass diffusion coefficient, , and mass and heat convective coefficients, and , thereby requiring a numerical solution. The finite volume method was used in this work owing to its local conservative character and discretisation/implementation simplicity. The space derivatives are discretized by central second-order formulae whereas the time derivative is approached by the second-order accurate Cranck-Nicholson scheme. The resulting systems of algebraic equations are solved by the well-known TDMA algorithm. The procedure is classical in the literature and for more details the reader is referred to [19]. The numerical approximations for the local problems were validated against analytical solutions for a one-dimensional plate subjected to convective transfer on both sides [9]. The largest local errors for the nondimensional parameters were smaller than 1.8% (very early in the process), substantially decreasing as the process advances. Further discussions and validation of the local approximation are presented in [12].

4. Numerical Examples

4.1. Validation of the Flow Rate Model

The present system approach combines a global solution to obtain flow rates and velocities inside the channels and a local approximation to heat and mass transfer for timber layers. The critical aspect of the present model is the accurate determination of the channel velocities across the timber stack. Validation is, therefore, performed for the flow solution against a CFD simulation of a reduced size dry kiln using the Ansys CFX commercial package.

The simulations are performed for two plenum sizes, 0.090 and 0.180 m, and 12 timber layers. The entrance velocity at the inlet plenum is assumed m/s for both cases. Table 1 indicates the geometrical data and flow information. It is important to note that, in this example, only the hydrodynamic solution is determined, especially the channel velocities across the stack height.

tab1
Table 1: Dimensions and operational parameters for a reduced size dry kiln.

The velocity distribution for plenum sizes (a) 0.090 m and (b) 0.180 m is presented in Figure 5. The same entrance velocity at the inlet plenum (located at in Figure 2) used in both cases gives rise to substantially different flow rates across the timber stack. In case (a), the total flow rate is m3/s, whereas the larger plenum of case (b) leads to a higher flow rate, m3/s. Since the stack height, lumber thickness, and channel height are the same for both settings, case (b) provides channel velocities twice as high as case (a); that is, m/s and m/s, respectively. Figure 5 shows that differences between the Ansys CFX and present approach are acceptably small, with average values 3.62% and 1.64% for cases (a) and (b), respectively.

538931.fig.005
Figure 5: Channel velocities across the stack height for plenum velocity m/s and plenum sizes (a) = 0.090 m and (b) 0.180 m.
4.2. Full Scale Simulation of a Dry Kiln

It has been well established in the literature that the air temperature and velocity dictate the drying rate. Therefore, the ideal compact dry kiln would have to combine not only uniform drying rates, but also high lumber volume and compact dimensions (e.g., by designing narrow inlet and exit plenums). Within this framework, this example focuses on two issues: (a) the aerodynamic behaviour of the dry kiln and (b) evolution of the lumber moisture content and drying rates.

The timber stack comprises 30 lumber layers () of thickness  m, as depicted in Figures 1 and 2. The other dry kiln dimensions and process parameters are presented in Table 2. The simulations were performed for plenum sizes 0.10, 0.15, 0.20, 0.30, 0.50, and 0.80 m. The fan performance curve is described using a fourth-order polynomial in of equation (9) with a coefficient of determination 0.99987, as indicated in Table 3. In the present case, the operating region is located after the instability zone, thereby ensuring smooth operating conditions.

tab2
Table 2: Dry kiln dimensions and operational parameters.
tab3
Table 3: Fan performance curve for a rotational speed  rpm: fourth order polynomial regression, , in of equation (9).
4.2.1. Hydrodynamic Solution: Velocity Distribution across the Timber Stack

This section highlights the effect of the plenum size on the hydrodynamic performance of the compact dry kiln using the formulation discussed in the previous sections. The system approach provides the operating point corresponding to the intersection between the fan performance curve and the system characteristic curve. Figure 6 presents the operating conditions for different plenum sizes, in which is the global flow rate and is the pressure difference between the entrance of the inlet plenum and exit of the exit plenum. Narrow plenums (e.g., in this example, 0.10 m) impose significant head losses leading to smaller global flow rates and higher pressure drops. As the plenum size increases, the wall effect decreases, permitting balance flow conditions at smaller pressure drops and higher flow rates, converging to m3/s and mm H2O for the present compact kiln. Theoretically, when and , the pressure in the inlet and exit plenums is uniform and is due only to the pressure drop in the flow channels (flow between two consecutive lumber layers) across the timber stack.

538931.fig.006
Figure 6: System approach: fan and system characteristic curves and operating points for plenum sizes = 0.10, 0.15, 0.20, 0.30, 0.50, and 0.80 m.

The operating point represents a global measure of the hydrodynamic balance in the dry kiln. For each such point depicted in Figure 6, the global flow rate, , encompasses individual contributions by each flow channel, as illustrated in Figures 3(a) and 3(b). As discussed in the previous paragraph, the dimensions of the inlet and exit plenums affect head losses, which in turn cause different flow rates and, consequently, air velocities, , across the timber stack. Therefore, narrow plenums entail higher head losses, causing air to flow primarily across the top channels and producing highly inhomogeneous flow rates and velocity distributions across the timber stack. Conversely, the virtually uniform pressure distributions in large plenums not only lead to higher global flow rates, but also make it possible uniform velocity distributions across the timber stack (highly desirable to ensure uniform drying rates). Figure 7 shows the velocity distribution in all 31 channels across the lumber stack, in which the differences caused by narrow plenums can be clearly observed. The simulations show that the relative differences between the maximum and minimum velocities increase from ~1% to ~77% for plenum sizes and , respectively.

538931.fig.007
Figure 7: Velocity distribution across the timber stack for plenum sizes 0.10, 0.15, 0.20, 0.30, 0.50, and 0.80 m.
4.2.2. Lumber Moisture Content and Drying Rates

The physical model dictated by the governing equations and corresponding boundary conditions establishes that, in general, higher velocities yield higher convective and mass transfer coefficients at the lumber-air interface. In addition, the strong temperature dependency of the moisture diffusivity leads to an equally strong correlation between the air flow temperature and moisture content evolution along the drying process. Therefore, a physical understanding of the dry kiln behaviour regarding both effects is essential to defining drying strategies within practical industrial applications. This section is focused on such physical aspects, in which the effects of the air flow temperature and plenum size are investigated. The timber and other kiln/process parameters are defined in Table 2. The time step used in the simulations to solve the heat and mass conservation equations is .

The previous section features the effect of the plenum size upon the velocity distribution across the timber stack. The ensuing effect upon the drying process for a narrow plenum is shown in Figures 8 and 9 which presents the average moisture content, , and the moisture content at (centre of the timber layer) for timber layers 1 (bottom of the stack) and 30 (top of the stack) corresponding to air flow temperatures 60°C, 90°C, and 120°C (333 K, 363 K, and 393 K). One could highlight two aspects regarding higher air flow temperatures: (i) drying rates increase and (ii) the effects of an inhomogeneous velocity distribution across the timber stack are magnified. Higher air flow temperatures cause the timber temperature and moisture diffusivity to increase, as (15) indicates, leading to higher moisture flow towards the timber-air interface with consequent increase in drying rates. Such higher moisture flow amplifies the sensitivity of the convective heat and mass transfer coefficients with regard to local velocities, leading in turn to greater differences of the drying rate across the timber stack. For instance, Figure 8 shows that the differences between the average moisture content for the 1st and 30th timber layers after 8 hours of drying time increases from 2% to 15% for air flow temperatures ranging from 60°C to 120°C (333 K to 393 K).

538931.fig.008
Figure 8: Average moisture content for a plenum 0.1 m.
538931.fig.009
Figure 9: Moisture content at for a plenum 0.1 m.

The heat and mass transfer coupled effects are featured in Figure 10 which shows the differences in heating and drying rates for the 1st and 30th timber layers for a narrow plenum. The aforementioned higher influence of the air flow temperature is expressed in higher heating and drying rates, that is, variations of the average temperature and moisture content within a single timber layer. Furthermore, the differences in air velocities across the timber stack (see Figure 7) cause substantial changes in heating and drying rates, especially early in the process.

538931.fig.0010
Figure 10: Heating rate, , versus dehumidification rate, , for air flow temperatures 60°C (333 K) and 120°C (393 K).

It is worthy to notice that, as the plenum size increases, differences in air velocity decrease, leading to homogeneous heating and drying rates across the timber stack. Figure 11 presents for the 1st and 30th timber layers and plenum sizes 0.1, 0.2, and 0.8 m, in which iso-time curves are also shown. The larger thermal diffusivity causes timber layers to heat at higher rates than the corresponding dehumidification rates, that is, after 2 hours of drying time all timber layers reach approximately the air flow temperature, whereas drying rates, , are still at higher levels. Figure 11 also shows that, for larger plenums, for example, 0.8 m, heating and drying rates across the timber stack are virtually uniform (the curves for 1st and 30th layers are almost coincident in the figure).

538931.fig.0011
Figure 11: Heating rate, , versus dehumidification rate, , for air flow temperature 120°C (393 K) and plenum sizes 0.1, 0.2, and 0.8 m.

The dynamics of the drying process is summarised in Figure 12 which shows for larger plenums (uniform heating and drying rates across the timber stack) and air flow temperatures 60°C, 90°C, and 120°C (333 K, 363 K, and 393 K). For the sake of greater accuracy, the iso-time curves are computed using intermediate air flow temperatures in the range between 60°C and 120°C. Firstly, as discussed early in this section, it is possible to observe that higher air flow temperatures lead to higher drying rates; for example, after 24 hours of drying time the average moisture contents are 0.16, 1.50, and 4.62 kg/kg for air flow temperatures 120°C, 90°C, and 60°C, respectively. It is also interesting to notice that, in all cases, the timber layers quickly reach temperature equilibrium with air flow (around 4 hours of drying time).

538931.fig.0012
Figure 12: Evolution of the average timber temperature, , and moisture content, , for plenum 0.8 m.

5. Conclusions

Convective drying is one of the most widely used drying strategies in industry. The process is performed in dry kilns, in which heated air flows through a lumber stack comprising timber layers separated by stickers. The physics of the problem encompasses the hydrodynamic description of the air flow, conjugated heat, and moisture transfer within timber layers and convective heat and moisture transfer at lumber-air interface. In this work, the actual air flow distribution is determined using a system approach, which provides the kiln operating flow condition (the point at which the fan performance curve at a given rotational speed intersects the system characteristic curve). The air flow velocities across the timber stack at the operating point are used to compute the convective heat and moisture coefficients at the timber surface, which in turn are used to solve the fully coupled heat and moisture transfer problems.

This class of problems requires definition of geometric and operational dry kiln parameters able to provide both high drying rates and uniform moisture contents across the timber stack. The former is justified by economic reasons, whereas the latter is required with the objective of ensuring good quality of the final product (avoid warping, decolouration, etc.). The most critical aspect of this problem is the correct computation of the channel velocities across the stack height. Therefore, in order to ensure that the present approach is acceptable, channel velocity solutions for a reduced size dry kiln are compared with the Ansys CFX commercial code. With the objective of evaluating the uniformity of the drying process across the stack height, the effects of the kiln geometric configuration (plenum sizes) and air flow temperatures are investigated. The fully coupled simulations of a compact dry kiln have shown that higher air flow temperatures increase drying rates (a positive aspect); however, the effects of a nonuniform velocity distribution are magnified, that is, inhomogeneous drying rates across the timber stack (a negative aspect). Therefore, an industrial solution must seek a compromise between both effects, which may eventually lead to changing the sticker number and dimensions (to improve velocity uniformity), establish progressive air heating (to reduce the effect of inhomogeneous velocity distributions) and even replace the air circulating system (providing a different fan characteristic curve).

Nomenclature

Area (m2)
:Specific heat (J kg−1 K−1)
:Diffusion coefficient of the component A in B (m2 s−1)
:Hydraulic diameter (m)
:Diffusion coefficient of moisture in the timber (m2 s−1)
:Timber roughness (m)
:Fan characteristic equation
:System characteristic equation
:Convective heat transfer coefficient (W m−2 K−1)
:Convective mass transfer coefficient (m s−1)
:Stack height (m)
:Sticker height (m)
:Lumber thickness (m)
:Stack level
:Loss coefficient
:Flow channel length (m)
:Stack length (m)
:Sticker head width (m)
:Individual channel length (m)
:Lewis number
:Moisture content (kg of moisture/kg of dry wood)
:Order of the polynomial-fan equation
:Order of the polynomial-system equation
:Vertical number of flow channels across the stack height
:Number of timber layers across the stack height
:Number of stickers between two timber layers
:Number of flow channels between two timber layers
:Average velocity (m s−1)
:Pressure (mmH2O)
:Pressure in the flow channel (mmH2O)
:Prandtl number
:Flow rate (m3 s−1)
:Hydraulic resistance (Pa (m−3 s−1)−2)
Re:Reynolds number
:Temperature (°C)
:Stack width (m)
:Maximum inlet/exit plenum width (m)
:Minimum inlet/exit plenum width (m)
:Vertical axis (m).
Greek Letters
:Kinetic energy correction factor
:Thermal diffusivity of the timber (m2 s−1)
:Polynomial coefficient-fan equation
:Polynomial coefficient-system equation
:Heat conductivity (W m−1 K−1)
:Friction coefficient
:Viscosity (Pa s)
:Specific mass (kg m−3).
Subscripts
:Air
:Operating point
:Stack level
:Plenum
:Flow between two timber layers
:Timber
:Interface air-timber
:Flow channel entrance
:Flow channel exit.
Superscripts
:Total measure between two timber layers
:Inlet plenum
:Parallel channels between stickers
:Exit plenum
:Plenum
:Flow between two timber layers.

Conflict of Interests

The authors affirm that they have no financial support or conflict of interests with manufacturers of any commercial software mentioned in the paper.

References

  1. R. Shmulsky and P. D. Jones, Forest Products & Wood Science, Wiley, Chichester, UK, 6th edition, 2011.
  2. V. P. C. Mohan and P. Talukdar, “Three dimensional numerical modeling of simultaneous heat and moisture transfer in a moist object subjected to convective drying,” International Journal of Heat and Mass Transfer, vol. 53, no. 21-22, pp. 4638–4650, 2010. View at Publisher · View at Google Scholar · View at Scopus
  3. M. Riepen and B. Paarhuis, “Analysis and optimisation of the airflow distribution in convection kilns,” in Proceedings of the 1st Workshop on State of the Art for Kiln Drying, Edinburgh, UK, October 1999.
  4. A. Kaya, O. Aydin, and I. Dincer, “Numerical modeling of heat and mass transfer during forced convection drying of rectangular moist objects,” International Journal of Heat and Mass Transfer, vol. 49, no. 17-18, pp. 3094–3103, 2006. View at Publisher · View at Google Scholar · View at Scopus
  5. A. Kaya, O. Aydin, and I. Dincer, “Numerical modeling of forced-convection drying of cylindrical moist objects,” Numerical Heat Transfer Part A, vol. 51, no. 9, pp. 843–854, 2007. View at Publisher · View at Google Scholar · View at Scopus
  6. A. Kaya, O. Aydin, and I. Dincer, “Heat and mass transfer modeling of recirculating flows during air drying of moist objects for various dryer configurations,” Numerical Heat Transfer Part A, vol. 53, no. 1, pp. 18–34, 2008. View at Publisher · View at Google Scholar · View at Scopus
  7. R. Younsi, D. Kocaefe, S. Poncsak, and Y. Kocaefe, “Computational and experimental analysis of high temperature thermal treatment of wood based on ThermoWood technology,” International Communications in Heat and Mass Transfer, vol. 37, no. 1, pp. 21–28, 2010. View at Publisher · View at Google Scholar · View at Scopus
  8. F. Thibeault, D. Marceau, R. Younsi, and D. Kocaefe, “Numerical and experimental validation of thermo-hygro-mechanical behaviour of wood during drying process,” International Communications in Heat and Mass Transfer, vol. 37, no. 7, pp. 756–760, 2010. View at Publisher · View at Google Scholar · View at Scopus
  9. C. Hirsch, Numerical Computation of Internal and External Flows: The Fundamentals of Computational Fluid Dynamics, vol. 1, Butterworth-Heinemann, Oxford, UK, 2nd edition, 2007.
  10. P. S. B. Zdanski, M. A. Ortega, and N. G. C. R. Fico Jr., “Heat transfer studies in the flow over shallow cavities,” Journal of Heat Transfer, vol. 127, no. 7, pp. 699–712, 2005. View at Publisher · View at Google Scholar · View at Scopus
  11. P. S. B. Zdanski, M. Vaz, and A. P. C. Dias, “Forced convection heat transfer of polymer melt flow inside channels with contraction/expansion sections,” International Communications in Heat and Mass Transfer, vol. 38, no. 10, pp. 1335–1339, 2011. View at Publisher · View at Google Scholar · View at Scopus
  12. P. S. B. Zdanski, M. Vaz Jr., R. F. L. Cerqueira et al., “Numerical simulation of the coupling between transient heat and mass diffusion process,” in Proceedings of the 7th National Congress of Mechanical Engineering, São Luis, Brazil, July 2012, (Portuguese).
  13. J. J. Nijdam and R. B. Keey, “The influence of kiln geometry on flow maldistribution across timber stacks in kilns,” Drying Technology, vol. 18, no. 8, pp. 1865–1877, 2000. View at Scopus
  14. S. Post, Applied Computational Fluid Mechanics, Jones and Barlett, Sudbury, Mass, USA, 2011.
  15. P. K. Swamee and A. K. Jain, “Explicit equations for pipe-flow problems,” Journal of the Hydraulics Division, vol. 102, no. 5, pp. 657–664, 1976. View at Scopus
  16. A. Thumann and D. Mehta, Handbook of Energy Engineering, Fairmont Press, Lilburn, Ga, USA, 6th edition, 2008.
  17. H. E. Musch, G. W. Barton, T. A. G. Langrish, and A. S. Brooke, “Nonlinear model predictive control of timber drying,” Computers and Chemical Engineering, vol. 22, no. 3, pp. 415–425, 1998. View at Scopus
  18. T. L. Bergman, A. S. Lavine, F. P. Incropera, and D. P. DeWitt, Fundamentals of Heat and Mass Transfer, Wiley, Chichester, UK, 7th edition, 2011.
  19. S. V. Patankar, Numerical Heat Transfer and Fluid Flow, Hemisphere, Washington, DC, USA, 1980.