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Dynamic modeling of waste incineration plants with rotary kilns

1998, Chemical Engineering Science

Incineration is not a new technology and has been used to destroy organic hazardous waste for many years. Nevertheless, variability in waste composition and the severity of the incineration operating conditions may result in many practical operating problems, high maintenance requirements and equipment unreliability. Moreover, a large number of constraints must be satisfied. These constraints are imposed by permit, by design and by operating practice. The goal of this work is to present a dynamic model which is able to follow large variations in process conditions and to be of practical value from a control point of view. This paper deals in particular with the development of a dynamic model which describes the behavior of a rotary kiln (primary combustion chamber with heterogeneous combustion) as well as the corresponding afterburner system (secondary combustion chamber with homogeneous combustion) followed by a heat recovery system which completes the hot section of a typical incineration plant. Special attention was devoted to a start-up procedure which was used as a specific application, to check the robustness and reliability of the model itself effectively. In addition, a number of comparisons with experimental data available from commercial units are reported to complete the model validation under ordinary feeding conditions.

PII: Chemical Engineering Science, Vol. 53, No. 15, pp. 2727—2742, 1998 ( 1998 Elsevier Science Ltd. All rights reserved Printed in Great Britain S0009–2509(98)00081–5 0009—2509/98/$—See front matter Dynamic modeling of waste incineration plants with rotary kilns: Comparisons between experimental and simulation data M. Rovaglio,* D. Manca and G. Biardi Dipartimento di Chimica Industriale ed Ingegneria Chimica ‘‘G. Natta’’, Politecnico di Milano, Italy (Received 4 September 1996; accepted 31 January 1998) Abstract—Incineration is not a new technology and has been used to destroy organic hazardous waste for many years. Nevertheless, variability in waste composition and the severity of the incineration operating conditions may result in many practical operating problems, high maintenance requirements and equipment unreliability. Moreover, a large number of constraints must be satisfied. These constraints are imposed by permit, by design and by operating practice. The goal of this work is to present a dynamic model which is able to follow large variations in process conditions and to be of practical value from a control point of view. This paper deals in particular with the development of a dynamic model which describes the behavior of a rotary kiln (primary combustion chamber with heterogeneous combustion) as well as the corresponding afterburner system (secondary combustion chamber with homogeneous combustion) followed by a heat recovery system which completes the hot section of a typical incineration plant. Special attention was devoted to a start-up procedure which was used as a specific application, to check the robustness and reliability of the model itself effectively. In addition, a number of comparisons with experimental data available from commercial units are reported to complete the model validation under ordinary feeding conditions. ( 1998 Elsevier Science Ltd. All rights reserved. Keywords: Incineration; dynamics; modeling; combustion; control. INTRODUCTION Incineration processes are today considered important components of waste management policy all over the world. The next several decades should see an increase in the demand for incineration because of recent limitations placed on the remaining hazardous waste management options. Waste incineration offers the following real or at least potential advantages: f volume reduction, especially for solids with a high combustible content; f detoxification, especially for combustible carcinogens, pathologically contaminated material, toxic organic compounds, etc.; f environmental impact mitigation, by destruction of all the undesired secondary effluents or byproducts which would create further significant pollution problems; * Corresponding author. Fax: 00 39 02 70 63 81 73; e-mail: [email protected]. f energy recovery, especially when large quantities of waste are available and users of heat or steam are located nearby. These advantages have justified the development of a variety of incineration systems showing greatly varying degrees of complexity and process schemes to meet the needs of municipalities, commercial and industrial firms and institutions. The most commonly used incineration techniques include off-gas treatments, liquid injection and rotary kilns. Although off-gas and liquid injection incinerators have been the most popular in Europe and the United States, rotary kiln incinerators are better suited to handle all physical forms of hazardous wastes and new high temperature chambers are capable of destroying highly toxic organic wastes (e.g. chlorinated organic compounds, PCBs) very efficiently. Multiple hearth, fluidized bed and other technologies such as cement kiln and molten salt combustion seem to be promising but they need further demonstration to prove that they are economically, technically and environmentally acceptable (Theodore and Reynolds, 1987). 2727 2728 M. Rovaglio et al. Incineration is not a new technology and has been used for treating organic hazardous waste for many years. However, variability in waste composition and the severity of the incineration operating conditions may result in many practical operating problems, high maintenance requirements and equipment unreliability. In practice, the majority of incinerators are on-site facilities responsible for destroying a well-defined set of wastes. These incinerators must be robust enough to handle large variations in waste composition and heat release. In contrast, commercial facilities receive a wide variety of wastes from many generators. These wastes can vary in physical state (solid, liquid, sludge), chemical composition and thermodynamic properties. For both commercial and on-site incinerators, there is a large number of operating constraints. These constraints are imposed by permit, by design and by operating practices. Thus, the operator must blend the waste quality and control the waste flowrates, air flowrates, water scrubbing flowrates and auxiliary fuel flowrates such that all the constraints are satisfied. Moreover, operators seek to maximize the waste throughput or minimize the auxiliary fuel flow while maintaining a correct oxygen concentration value in the outlet gas. Incinerator operators currently attempt to achieve these objectives by intuition and experience (Behmanesh et al., 1990). Finally, feedstocks to incinerators can be continuous, semi-continuous or batch-wise which means that the time evolution of process conditions seems to be an important element in evaluating feasible and reliable operations. The goal of this work is to present a dynamic model being able to follow and describe large variations in terms of process conditions while being of real practical value from a control point of view. In the following presentation, the paper will deal with the development of a dynamic model capable of describing the behavior of a rotary kiln (primary combustion chamber with heterogeneous combustion) and the corresponding afterburner system (secondary combustion chamber with homogeneous combustion) plus the heat recovery system which completes the hot section of a typical incineration plant. Special attention was devoted to a start-up procedure which was used as a specific application, to check the robustness and reliability of the model itself effectively. Several comparisons with experimental data available from commercial units are also reported to complete the model validation under normal feeding conditions. INCINERATOR MODEL DEVELOPMENT The basic structure of the plant analyzed in this paper comprises, as mentioned above, a rotary kiln followed by a postcombustion chamber (afterburner system) and the corresponding air pre-heaters which allow the heterogeneous combustion to be improved and the auxiliary fuel consumption to be reduced. This configuration with geometry and plant layout is shown in Fig. 1. The solid is fed into the rotary kiln where the bulk of the combustion takes place and then moves along the chamber in a countercurrent configuration with respect to the gas flow. The resulting combustion products go to a cyclone postcombustion chamber which allows the total combustion and destruction of contaminants be completed. At the end of the plant’s hot section two spiral heat exchangers are used to preheat the incoming combustion air. Rotary kilns involve a large variety of combinations of processes such as particulate mixing, gas—solid or solid-phase reactions with intensive heat and mass transfer. The kiln analyzed here is operated with a semicontinuous feed and continuous discharge, pitched slightly from inlet to outlet and normally running less than half-full of solid. The upper space accommodates a countercurrent flow of hot gas which may serve both to supply the heat to the inlet stream and to carry out gaseous products. As mentioned above, the rotary kiln has a secondary chamber to ensure complete combustion of the waste. In practice, the kiln acts as the primary chamber to volatilize and partially or totally oxidize the combustible materials in the waste. Inert ash is then removed from the lower end of the kiln. The volatilized combustibles exit the kiln and enter the secondary chamber where Fig. 1. The hot section of an incineration process with combustion air preheating. Dynamic modeling of waste incineration plants additional oxygen is made available and ignitable liquid wastes or fuel can be introduced to achieve the desired operating temperature. In order to describe the time evolution of a combustion process like this, a simplified but rather general model can be built up on the basis of the following five main steps. ¼aste feedstock characterization and combustion stoichiometry The feedstock consists of a dry solid fraction (1!u ) and a moisture fraction u (mass fracH2O H2O tions). Within the dry solid, it is then necessary to distinguish between a combustible or burning fraction u which gives rise to gas products, and an inert BF fraction representing what will be also the final solid ash amount. A formal stoichiometry must be associated with the combustible fraction of the solid waste in order to account for the heterogeneous combustion. This leads to the following stoichiometric relationship: n!y C H O S N Cl #l O P mCO# m n p q x y O2 2 2 x!a ]H O#qSO # N #aNO#yHCl. 2 2 2 2 It must be emphasized that, as in the case of the heterogeneous combustion, CO has been assumed to be the main oxidation product of carbon. The total combustion to CO takes place in the gas phase. 2 Furthermore, waste compounds containing nitrogen are converted to N and NO according to the scheme 2 and the relationship reported by Bowman (1975) which evaluates the a fraction as a function of kiln temperature and oxygen excess. The stoichiometric coefficients are likely to occur in the gas phase and they will be taken into account further on in the paper. 2. Solid waste mass balance The solid waste undergoing the heterogeneous combustion passes through the rotary kiln incinerator for a certain contact, or retention time, while the effect of the mechanical agitation induced by rotation consists of the renewal of the surface area exposed to hot gases containing oxygen. Here, on this area, both volatilization and combustion processes take place at a rate which is largely dominated by the oxygen transfer rate towards the surface since the temperature is high enough to reduce the importance of the chemical reactions kinetics. This is why in the model scheme, the rate of destruction of the waste combustible fraction was computed according to the oxygen mass transfer from the bulk gas phase to the solid flame surface where its molar fraction disappears as it is consumed, also in transient conditions, at the rate it diffuses. The bulk solid phase movement along the rotary kiln is not different in nature from that of similar units. The proper correlations of those cases and available in literature may be used for the case in 2729 hand. This applies for the retention time as reported by Freeman (1989): 1.77JbF¸ h" . SDN (1) Whence the outlet solid flowrate ¼S can be deter065 mined by linking it to the aforementioned retention time and to the instantaneous solid mass hold-up within the incinerator: M SDN ¼S " 40-*$"M . 065 40-*$ 1.77JbF¸ h (2) This correlation means that despite the transient situation, where the solid movement in the rotary kiln is concerned, the same relationships as in the steady state hold true using the instantaneous value of the variable M . In other words, the transient condi40-*$ tions related to both gas and solid flows are assumed to be fast with respect to mass and energy dynamic balances. In agreement with the aforementioned assumptions, the mass balance for the solid phase reads as follows: dM 40-*$"¼ (1!u )!R !¼S . 8!45% H2O dr 065 dt (3) Particular attention must be devoted to the evaluation of R which represents the burned fraction of dr the solid waste. This term is strictly related to the transport phenomena active at the solid surface. The top plane of the solid phase is evidently the most likely region for the highest rates of both convective and radiative heat transfer as well as mass transfer to the contiguous gas phase (Ferron and Singh, 1991). This source term, therefore must be calculated in agreement with the scheme previously mentioned and is represented as follows: K (X"6-,!X*/ )PM C O2 O2 505 A. (4) R " x O2 dr k O2 The evaluation of A, the interface area for solid oxygen contact, would in principle require the knowledge of both gas-and solid phase motions inside the rotary kiln. This area is a function of the kiln geometry and the operating conditions. The maximum possible interface area would depend also on particle diameter and solid angle of repose; if one could estimate the average renewal surface velocity [A (m2/s)] then r A"A h. (5) r Jacob and Perlmutter (1980) have already provided criteria for determining such a renewal surface velocity. Nevertheless, we thought it more reliable and practical to treat A as an adaptive parameter based on the interpolation of a data matrix obtained from steady-state simulations of known kiln performances. As a matter of fact, if a given commercial unit can be operated in a specified range of operating conditions with satisfactory performances with respect to 2730 M. Rovaglio et al. waste residence time and combustion efficiency, the interface area can be easily derived from steady-state simulation as a function of the mass holdup and available oxygen mole fraction by inverting relation (4), and by noting that, at the steady-state conditions, the rate of destruction of waste combustible fraction is closely related to such parameters. From the solidphase balance, eq. (3), one gets ¼ (1!u )!R !¼S "0. 8!45% H2O dr 065 (6) Under the assumption of complete waste destruction: ¼S "¼ (1!u )(1!u ) BF 065 8!45% H2O with u BF (7) the waste burning fraction, R follows from dr u BF R "¼ (1!u )u "¼S dr 8!45% H2O BF 0651!u M u BF " 40-*$ h 1!u Fig. 2. Flame area as a function of solid holdup and oxygen mole fraction inside kiln. BF (8) BF and finally 1M u k BF " 40-*$ O2 A" h *x K C PM 1!u x 505 O2 BF O2 (9) In agreement with the assumption already made about the fast dynamics of solid flows within the kiln, the same effective area is held under transient conditions. On the basis of experimental data available and through the use of a commercial package (CYCOM 6.0), an example of interface area evaluation is reported in Fig. 2. It can be observed that if the total mass holdup inside the kiln increases, the flame surface achieves a larger value with a variation that is a function of the bulk oxygen concentration. A small O mole fraction in the gas phase will obviously 2 require an increase in the surface area for a given mass holdup and kiln volume in order to achieve the total burning of the waste load (slow firing conditions). On the other hand, a large oxygen excess corresponds to a practically constant required interface area (high firing rate). Figure 2 also shows a typical domain where the kiln under examination operates. Comparing the value obtained by the interpolated surface (as a function of kiln conditions) with the result of Jacob and Perlmutter’s relationship, the interphase reaction area is defined as the maximum between these numbers. Gas-phase mass balances. Taking into account possible auxiliary fuel, if required, the total gas-phase mass balance, which applies to both the rotary kiln and postcombustion chamber, becomes dM '!4"¼*/ w!*3#¼*/ w!*3#(1!k )R !*3 N2 !*3 O2 O2 dr dt w #¼ !¼g #¼ 065 8!45% H2O CH4 (10) while the corresponding mass balances for each component can be specified by dM ¼' i"¼*/! 065 y #R » (11) i i 505 dt + PM y i i i i with i"CO, CO , H O, SO , N , HCl, NO, O , 2 2 2 2 2 where y "M /+ M is the gas mole fraction of i i i i species i. The flows of all the i species forming ¼*/ and the i reaction rates R can be determined with reference to i the following simplified reaction scheme and assumptions (Niessen, 1978): `O2 mCO#(n!y) H O (a) C H O S N Cl &" m n p q x y 2 2 (x!a) N #aNO#yHCl. #qSO # 2 2 2 (b) CH #O P CO #2H O. 4 2 2 2 (c) CH #3 O P CO#2H O. 4 2 2 2 (d) N #O b 2NO. 2 2 (e) CO#1 O P CO . 2 2 2 (g) CO#H ObCO #H . 2 2 2 where reaction (a) represents the previously mentioned overall waste destroying reaction producing CO as result of heterogeneous combustion. CO will be immediately converted to CO , should enough 2 oxygen be present to accomplish reaction (e), or otherwise following reaction (f ) when the oxygen amount is too small. Waste nitrogen is converted to N and NO. 2 The corresponding fractions are determined through a relationship reported by Bowman (1975) as a function of kiln temperature and oxygen excess. The equilibrium conditions determined by reaction (d) refer to the afterburner system (homogeneous combustion). If auxiliary fuel is needed to maintain the required temperature (e.g. postcombustion chamber), reactions (b) and (c) allow total or partial methane combustion as a function of the oxygen availability to be taken into Dynamic modeling of waste incineration plants account. Reactions (a)—(c) and (e) are very fast and can be considered under equilibrium conditions (completely displaced towards CO and CO ) within the 2 range of temperatures usually adopted (900—1200°C). The extent of conversion for such reactions is in practice determined by the oxygen excess. Only reactions (d) and (f ) have been considered in a kinetic regime, while the corresponding rate expressions have been deduced by generalizing the theoretical analysis reported by Westenberg (1971). In particular, by considering the kiln, burner and postcombustion chamber residence times separately, it is possible to evaluate, at least qualitatively, the contributions to the production of NO corresponding to the different portions of the plant. Finally, the outlet flue gas flowrate ¼' re065 ported in eq. (11) can be determined on the basis of a simple fluid dynamics description as shown later in the paper. Energy balances Inside incinerators the heat is generally transferred to the walls or to the upper surface of the solid beds (whenever present) by radiation and convection and to the lower (covered) surface by the regenerative action of the rotating kiln walls. The radiating gases in the kiln free volume (i.e. not occupied by the waste) are also present within the confines of the burner and/or waste flames so that they may be considered as covering the entire free volume. Thus, due to the high gas temperature, solids and exposed walls receive heat primarily by radiation from the gas volume while both convective and regenerative heat flow play only a minor role in the overall heat transfer process. On the basis of the following assumptions: f perfectly mixed conditions; f equilibrium condition between gas and solid phase (if present); f complete combustion (for both waste and auxiliary fuel); a zero-dimensional model (lumped model) describing the variation of bulk flow temperature with time can be derived as follows: dº 505"¼*/ H*/ #¼ Q !*3 !*3 8!45% 8!45% dt where #¼ Q !¼g H !sGS(¹4!¹4 ) CH4 CH4 065 065 ' w,*/ !h pD¸(¹ !¹ )!¼s CM (¹ !¹ ) 3*& */ g w,*/5 065 pS g (12) º "M C (¹ !¹ ) 505 40-*$ vS g 3*& M Tg i C (¹ ) d¹ #M + vi '!4 + M i i i T3*& T*/!*3 C (¹ ) d¹ H*/ " p!*3 !*3 T3*& M Tg i H "+ C (¹ ) d¹. 065 pi +M i i i T3*& P P P (13) (14) (15) 2731 Equation (12) implies that waste and fuel are fed at the reference temperature. GS can be evaluated by means of the ‘well-stirred combustion chamber’ model reported by Hottel-Sarofim (1968): PP GS" A K d» dS cos 0q(r) q , . (1/C e )#(1/e ) nr2 V S gp G (16) The evaluation of the gas emissivity e refers to the G diagram reported by Hottel and Sarofim which only considers the emission band related to the combustion products: water and carbon dioxide. The emissivity of such components can be calculated through a polynomial regression of the experimental data as a function of the sum of the single partial pressures (P "P #P ) and of the beam length (¸ ): CW CO2 H2O W 11 log (¹e )" + C xn~1 10 G n n/1 where x"log (¸ P ). 10 W CW Coefficients C are deduced by interpolating the n diagrams reported by Hottel and Sarofim (1968) and the resulting values are summarized in Table 1. The convective heat transfer coefficient is computed by a Dittus—Boelter relationship hD Nu"0.023Re0.8Pr1@3, Nu" k (17) and corrected to take into account possible entry effects (see Knudsen and Katz, 1958) A B D . h "h 1#1.4 in ¸ (18) As for the transient heat conduction within the furnace walls, the problem can be easily modeled by means of the classical equation of heat diffusion in one dimension to be solved for adjacent layers of different materials and with proper boundary conditions ensuring the continuity of heat fluxes (Biardi et al., 1987). The real problem lies in choosing which discretization method is to be adopted and which numerical algorithm to use for integration purposes. While the latter question is discussed in the next paragraph, it is worth saying a few words about the former one. Figure 3 Table 1. Polynomial coefficients for gas emissivity evaluation C " 9.695]10~3 11 C " 5.407]10~2 10 C " 1.349]10~1 9 C " 2.068]10~1 8 C " 1.496]10~1 7 C "!1.049]10~1 6 C "!2.239]10~1 5 C " 3.061]10~2 4 C " 7.615]10~2 3 C " 4.062]10~1 2 C " 2.44956 1 2732 M. Rovaglio et al. reported by Owens et al. (1991) where it is clearly shown that the bed temperature evolution along the kiln is quite fast and can be considered of relevant value only for waste with high moisture content (sludges) or for contaminants desorbing applications. Fluid dynamic relationships Q "c JP !P -%!, 1 %95 ,*-/ l2 P "P #c o 065,,*-/ ,*-/ 1045 2 2 Fig. 3. Refractory and insulating discretization scheme with nodal arrangement. shows the case of a chamber wall divided into several shells or layers of refractory and insulating materials (onion scheme). The nodal arrangement distinguishes between volume-averaged temperatures, whose index is an integer, and surface-averaged temperatures, whose index is arbitrarily set as an integer plus or minus a half. Symbolically, this gives rise to the following thermal balance: d¹W KW n" n A r Cp (¹ !¹ ) W W dt n *X/2 -/,n~1@2 n~1@2 KW (¹ !¹ ) ! n A n`1@2 *X/2 -/,n`1@2 n for n"1, N (19) (22) (23) l2 P "P #c o 065,1045 1045 14541*3!3 2 (24) l2 P "P #c o 065,41*3!- . 14541*3!l 4 2 (25) Equations (22)—(25) reflect the level of the fluid dynamic simplification used within the model. Equation (22), in particular, allows an evaluation of air leakage due to the vacuum conditions usually maintained inside a waste kiln and, when necessary, this term will be included in the corresponding mass and energy balances. Equations (23)—(25) define the kiln pressure (P ), the postcombustion pressure (P ) and the ,*-/ 1045 first spiral heat exchanger pressure (P ) as 14541*3!a function of the gas flow, the geometry and the pressure level at the final outlet throat (P ). Values for 1 c are calculated from correlations available in litera* ture, see, for example, Rovaglio et al. (1990), and Douglas et al. (1980). AIR HEATER MODEL DEVELOPMENT In the plant configuration analyzed here (see Fig. 1), there is a heat recovery section consisting of a sequence of spiral heat exchangers placed after the burning system. These devices have a double function, namely: with the boundary conditions: hot side KW i A (¹ !¹ )"h nD¸(¹ !¹ ) l */ g w,*/ *X/2 -/,1@2 1@2 #GSp(¹4!¹4 ) (20) g w,*/ where ¹ ,¹ is the internal chamber wall tem1@2 w,*/ perature. cold side KW N A (¹ !¹ )"h A (¹ !¹ ) N`1@2 e e w,065 !." *X/2 I/,N N (21) where ¹ ,¹ is the external chamber wall N`1@2 w,065 temperature, where the overall chamber shell is divided into N different layers of thickness *X; KW is n the thermal conductivity and A represents the log -/,n mean surface area. Finally, it is worthwhile explaining the assumption about the equilibrium condition between gas phase and solid bed. This leads to the same temperature for both phases equal to the outlet temperature since it has been also assumed to be a CSTR scheme. The reason for this choice can be related to the results f air combustion pre-heating to improve the thermal yield of the plant; f gas cooling to reduce water consumption inside the scrubbing section. These apparatuses consist of a metal shell which covers a cylinder of refractory and insulating material where, inside the enclosure, a coil of special alloy is placed. The cold air flows inside the coils while the hot gas moves along the refractory lined chamber. A general scheme for the examined plant including such heat exchangers together with a schematic representation of the reference geometry is reported in Fig. 4. A new set of equations is needed to solve this problem. Inside the heat exchangers, the hot gas temperature can be considered uniform, that is, a CSTR scheme is assumed to be a consequence of the high turbulence due to the gas velocity and to the presence of coils: dº 505"¼ H !¼ H !Q !Q f */ f 065 -044%4 %9 dt (26) Dynamic modeling of waste incineration plants 2733 Fig. 4. Reference scheme for combustion air pre-heaters. where H*/ and H065 can be evaluated through eq. (15) while ¼ represents the hot gas flowrate assumed to f be equal for inlet and outlet flows. The heat losses Q through the walls can be -044%4 determined by the correlations reported by Hottel and Sarofim (1968): C A B D 1!k3 h */ #A Q " GS 8!-- 1!k4 8!--,n 4p¹3 -044%4 f ]p(¹4!¹4 ) (27) f 8!--,n ¹ k" 8!-(28) ¹ f where the term inside the square brackets defines the global heat transfer coefficient which takes into account both the convective and the radiative terms. In a similar manner, the heat exchanged between smokes and coils can be determined on the basis of the following relationship: NS Q " + pGS (¹4 —¹4 )#h A (¹ !¹ ) %9 #0*-4 f 8!--,n e n f 8!--,n n/1 (29) where the external convective coefficient h can be e evaluated through the Nu number given by Nu"0.110Re0.675Pr1@3 (30) which refers to a system with a cross-flow geometry (see Incropera and DeWitt, 1988). In all the previous correlations the variables corresponding to the coils are indicated with a subscript n implying a discretized description along the coil as sketched in Fig. 5. This means that the temperature evolution along the coils can be evaluated through a system of N equations: s ¼ H !¼ H #Q "0 !*3 !*3,n~1 !*3 !*3,n %9,n with n"1,2 ,N (31) s where the energy holdup corresponding to each single element is considered negligible since, with respect to the space discretization, the air temperature can be assumed instantaneously to be under the steady-state conditions. Finally, the metal coil temperature can be determined on the basis of the heat fluxes congruence at the 2734 M. Rovaglio et al. the global differential system. In mathematical terms this means M o c (¹W!¹ ) '!4: w pw n 3*& . ¼ kW 065 1 A (¹ !¹ ) l *x/2 -/,1@2 1@2 Fig. 5. Coils discretization scheme. wall so that, for each element, the result is (¹4!¹4 )#h A (¹ !¹ ) pGS #0*-4,n f 8!--,n e n f 8!--,n "h A (¹ !¹ ) */ n 8!--,n !*3,n with n"1, 2 , N s (32) where the internal heat transfer coefficient h is com*/ puted by a conventional Dittus—Boelter relationship. (33) Practically, in the case examined here, N was determined to be in the range of 8—10. Since the overall set of equations constitutes a coupled system of algebraic and differential equations behaving as a numerically stiff system, due to the large spectrum of characteristic times involved, a completely implicit method able to automatically select the proper integration step was used in order to achieve a fast and robust numerical solution. The well known LSODE package (Hindmarsh, 1983) was adopted for these purposes. ALGORITHM The overall dynamic model consists of 2(N #N#2)#1 differential equations, namely: c 1 mass balance for the solid phase [eq. (3)]; 1 total mass balance for the gas phase [eq. (10)]; N !1 mass balances for N !1 single components of c c the mixture [eq. (11)]; 1 energy balance for the gas phase [eq. (12)]; N!energy balances for the refractory and insulating layers [eq. (19)]; 1 energy balance for the hot gas phase inside the heat exchanger [eq. (26)]. In addition, 2N #8 algebraic equations which can be s summarized as follows: 4 fluid dynamics relationships [eq. (22)—(25)]; 4 boundary conditions [eq. (20)—(21)]; N air thermal balances inside coil [eq. (31)]; s N heat fluxes equality [eq. (32)] s complete the structure of the model which must be numerically integrated to describe the dynamic evolution of the hot section of an incineration process. In particular N, the number of refractory-insulating elementary layers, can be estimated on the basis of the minimum spectrum width for the time constants of MODEL VALIDATION The results reported below refer to a specific commercial unit of 0.9—1.2 MW total burning capacity (waste plus auxiliary fuel). Geometry and characteristics are briefly summarized in Table 2. The model validation, as the title states, is unfortunately based only on a limited number of variables measured and available from the plant, namely: f outlet gas temperature from the rotary kiln, f outlet gas temperature from the postcombustion chamber, f external skin temperature for the rotary kiln, f external skin temperature for the postcombustion chamber, f oxygen mole fraction in the postcombustion gas outlet, f air temperatures from heat exchangers. However, the comparison is completed by the knowledge of the transient evolution of some input variables, such as: — waste flowrate to the rotary kiln, — fuel flowrates to kiln and postcombustion chamber, — air flowrate to the spiral heat exchangers. Table 2. Plant description Internal length (m) Internal diameter (m) Refractory thickness (m) Refractory conductivity (W/m K) Refractory heat capacity (J/kg K) Refractory density (kg/m3) Insulating thickness (m) Insulating conductivity (W/m K) Insulating heat capacity (J/kg K) Insulating density (kg/m3) Rotary kiln Postcombustion chamber 4.43 1.515 0.100 1.696 700.0 2550.0 0.120 0.29 500.0 890.0 3.664 1.722 0.110 1.348 600.0 2000.0 0.125 0.198 400.0 475.0 Dynamic modeling of waste incineration plants With reference to the plant measurements, it must be specified that the outlet gas temperatures are measured through a thermocouple placed in the outlet gas exit which has a small diameter with respect to the unit one. Moreover, both the kiln thermocouples are shielded from radiation by a special enclosure made in the refractory material. Therefore, the temperatures measured can be assumed to represent the average conditions of the outlet smokes. As a consequence of the assumed perfectly stirred conditions, defined by the turbulent regime, such temperatures also correspond to the average temperatures of the units. Finally, the reported skin temperatures derive from the average of several measurements taken with a manual thermocouple placed along the cylinder. Start-up procedure The start-up operation is a long transient condition requiring about one day to achieve the final values of the operating temperatures (1200°C for the afterburner system). A slow warm-up is necessary to avoid sudden expansion effects which may compromise the robustness and the reliability of refractory and insulating materials. Therefore, the burners, operating in the manual control mode, are brought to their maximum fuel capacities over a period of 2—3 h, with each starting at a different time. Only when the correct temperature regime is reached, the waste can be loaded and a high efficiency of total combustion can be achieved. Simulations and comparisons with experimental measurements taken during operation can be of real use in evaluating the reliability of the model’s basic structure without introducing possible misunderstandings due to results affected by wrong waste characterization in terms of feed flowrate, composition or heat of combustion. A schematic representation of the main input variables during the transient is reported in Fig. 6 referring to the first 11 h of the diagrams. Fuel and air inlet flowrates correspond to flows imposed by valves operating in the manual mode. The variation of such variables is modeled as ramp disturbances derived from the corresponding experimental measurements. The evolution of the postcombustion fuel flowrate also includes a perturbation related to a fuel-line blockage. Once again, with reference to the first 11 transient hours, a comparison between experimental data available and model evaluation as a function of time is reported in Figs 7 and 8 for the outlet gas temperature from the rotary kiln and the postcombustion chamber, respectively. Both transients show a very satisfactory agreement with the corresponding experimental data, not only in terms of characteristic times, but also in terms of absolute values and variations. A further model validation can be obtained by comparing the external kiln surface temperatures as reported in Figs 9 and 10. Although the number of available experimental measurements is very limited, the agreement between predictions and measurements confirms that 2735 significant information about the refractory and insulating temperature profiles can be derived from model simulations. One example of such profiles is reported in Fig. 11 as the evolution of temperature with time and refractory layers. The aforementioned analysis seems to confirm that the thermal capacities for both refractory-lined chambers and the flue gas together with the corresponding transport phenomena were well represented mathematically and that the related portion of the model can be considered reliable and useful for a further simulation study. ¼aste feeding conditions Waste characterization is a major factor in assessing the simulation of incineration processes by a deterministic model. The characterization of wastes is usually accomplished using the analytical techniques available although problems do arise since almost all the traditional analytical methods are only applicable for the analysis of either pure chemicals or nearly homogeneous materials. Wastes are typically highly heterogeneous and proper sample collecting and handling are obviously critical steps in determining waste characteristics and simulating the corresponding incineration processes. However, as it is shown below, reasonable estimates of combustion heat and ultimate analysis can be considered sufficient for model simulation purposes. All the simulation results reported here are refer to a waste characterization as given in Table 3. With reference to the previously described plant operation, waste feeding was commenced about 11 h after the initial startup, when the kiln temperature was around 850°C and the corresponding postcombustion chamber was at 950°C. At that moment the kiln burner was switched off and the postcombustion burner was put in the automatic operating mode to achieve an imposed setpoint of 1200°C. Combustion air flows and temperatures consequently followed the pattern described above. A schematic representation of the input variables evolution during the overall transient is shown in Fig. 6. The waste 40 l boxes weighting about 4 kg each are loaded by a manually actuated feeder at a rate of 1—2 boxes/minute. The average waste flowrate, corresponding to a period of 30 h, was estimated at about 340 kg/h. A global comparison between available experimental data and corresponding simulation data, including start-up and waste feeding conditions, is shown in Figs 7—12. It can be observed that the scattering of the experimental data of the kiln temperature, as shown in Fig. 7, can be explained on the basis of rapid changes in the waste feed rate as well as in the composition or heat of combustion. The corresponding simulation calculations refer to the average conditions mentioned earlier since the impact of the kiln fluctuations on the other variables seems to be very limited and this approximation does not affect the 2736 M. Rovaglio et al. Fig. 6. Input measured variables for start-up and loading conditions: (a) waste flowrate to the rotary kiln (b) fuel flowrates to rotary kiln and postcombustion chamber (c) air flowrate to the 1st spiral exchanger (d) air flowrate to the 2nd spiral exchanger. Dynamic modeling of waste incineration plants 2737 Fig. 7. Outlet gas temperature from Rotary kiln. Fig. 8. Outlet gas temperature from Postcombustion Chamber. model results. Figure 8 shows good agreement between the experimental postcombustion temperature and the one predicted by simulation. Transients of long duration, rapid variations of process variables and temperature control actions are all perfectly reproduced by the model. The satisfactory results can also be confirmed by observing the skin temperature variations for both of the refractory-lined chambers (see Figs 9 and 10): the typical reaction curves obtained coincide perfectly with the experimental data and validate not only the correlations adopted to describe heat fluxes but also the thermo-physical data used to characterize the corresponding insulating materials. Finally, Fig. 12 shows the time evolution of the oxygen mole fraction in the outlet section of the afterburner system. The comparison between experimental data and simulated ones shows an excellent agreement under the waste-feeding condition while the evident difference during the start-up procedure, indicates a failure of the oxygen sensor which shows an unrealistically high value of the measured quantity. In fact, since the auxiliary fuel consumption is known, a simple material balance clearly demonstrates that the total air flow needed to satisfy the experimentally indicated oxygen concentration is 3—4 times greater than the real maximum flow capacity of the plant itself. Conversely, the agreement shown for the wastefeeding conditions validates the model described from the material balance point of view. In particular, eq. (4), which was adopted to define the burning rate of the waste, the reaction scheme and the general characterization of the flows together with all the corresponding model assumptions, seems to be more than acceptable for the correct simulation of the hot section of an incineration plant. Once the model was proved reliable and capable of representing the phenomena involved, over a large range of operating conditions, some important information about unmeasurable variables could be easily deduced. An example is shown in Fig. 13 which shows the air 2738 M. Rovaglio et al. Fig. 9. External skin temperature for Rotary kiln. Fig. 10. External skin temperature for Postcombustion Chamber. leakage into the rotary kiln throughout the transient period. Knowledge and control of such a variable will certainly allow significant improvements to be taken in the auxiliary fuel consumption. Heat recovery Finally, Figs 14 and 15 show the comparison between the experimental and computed values of the outlet air temperature from the first and second heat exchanger respectively. In both cases there is a very good agreement between theory and practice. These last results, even though quite obvious on the basis of the previous ones, emphasize the reliability of the model itself since the presence of thermal and material recycles does not seem to compromise the robustness of the model simulation in object. CONCLUSIONS A generalized model for the dynamic simulation of incineration kilns was developed to reproduce the operating trends of some monitored variables for a commercial plant. Dynamic modeling of waste incineration plants 2739 Fig. 11. Refractory temperature profile versus time and grid index. Table 3. Waste characterization Component Mass fraction Net heat of combustion H O N C S Cl Inerts Water 0.0620 0.0900 0.0100 0.2840 0.0050 0.0080 0.1699 0.3711 8500 kJ/kg Fig. 12. Oxygen mole fraction in the outlet gas from postcombustion. The above analysis includes start-up procedures (considered as an important transient condition) together with more usual waste-feeding operations. All the results confirm that the model description can be considered to be generally applicable and easily extendable to different processes or plant situations where a reliable dynamic knowledge is available. Although the application examples of the aforementioned model have certainly been obtained on the basis of several minor assumptions and simplifying hypotheses, the close agreement between 2740 M. Rovaglio et al. Fig. 13. Air leakage for the Rotary Kiln. Fig. 14. Combustion air temperature from 1st heat exchanger. Fig. 15. Combustion air temperature from the 2nd heat exchanger. Dynamic modeling of waste incineration plants predicted values and experimental data clearly confirms that such a model can be of real value for engineering purposes. In the future, research will be devoted to exploiting the good quality of the simulated process behaviors in the application of the model to the definition of improved control strategies. In particular, a control structure, allowing stable and reliable operations, satisfying legal limits and process conditions and compatible with a maximum waste load, can be defined as a challenging issue for future study. NOTATION A A e A -/,n A n A r A q C p!*3 C pi CM pS C pw C vS C s C 505 D F GS h */ h e H*/ !*3 H */ H 065 K KW n K x k ¸ ¸ W M '!4 M i M 40-*$ N N s interface area, i.e. the flame surface, m2 external skin area, m2 log mean surface area at level n of the discretization grid, m2 wall area at section ‘n’ of the spiral exchanger, m2 average renewal surface velocity, m2/s total refractory surface, m2 air heat capacity at constant pressure, J/kg K gas heat capacity at constant pressure of component ‘i’, J/kg K mean solid heat capacity at constant pressure, J/kg K wall heat capacity at constant pressure, J/kg K solid heat capacity at constant volume, J/kg K ratio between sink surface and total surface,1 total concentration in gas phase, kmol/m3 internal combustion chamber diameter, m factor: 1 for undammed kilns;'1 for dammed kilns gas—wall surface exchange area, m2 internal convective heat transfer coefficient, W/m2 K external convective heat transfer coefficient, W/m2 K inlet air enthalpy, J/kg inlet gas enthalpy, J/kg outlet gas enthalpy, J/kg absorption coefficient, 1/m thermal conductivity at level ‘n’ of the discretization grid, W/m K mass transfer coefficient (m/s) evaluated on the basis of the following dimensionless relationship: Sh "0.023 Re0.8 Sc1@3 D D thermal conductivity, W/m K internal combustion chamber length, m free mean path of a photon inside the combustion chamber, ft gas holdup inside the kiln, kg molar holdup for component ‘i’ kmol solid holdup inside kiln, kg rotational speed of kiln, rps total number of discrete coil sections Nu P CW P %95 P ,*-/ PM i Pr Q %9 Q -%!, Q -044%4 Q CH4 Q 8!45% Re R dr R i S Sc Sh ¹065 a ¹*/ !*3 ¹*/ f ¹ f ¹065 f ¹ g ¹W n ¹ 3*& ¹ w,*/ ¹ 8!-¹ 8!--,n º 505 v 065 » 505 ¼065 a ¼*/ !*3 ¼ CH4 ¼*/ f ¼065 f ¼*/ i ¼g 065 ¼S 065 ¼ 8!45% X"6-, O2 X*/ O2 2741 Nusselt number partial pressure of CO and H O in the gas, 2 2 atm external pressure, Pa internal kiln pressure, Pa molecular weight of the component, ‘i’ (kg/kmol) Prandtl number heat exchanged between smokes and coils, W air leakage flowrate, m3/s heat losses through the exchanger walls, W methane net heat of combustion, J/kg waste net heat of combustion, J/kg Reynolds number destroying rate of waste combustible fraction, kg/s production/consumption for reactions in kinetic regime, kmol/m3 s kiln slope (degrees from horizontal) Schmidt number Sherwood number outlet air temperature, K inlet air temperature, K inlet flue gas temperature, K flue gas temperature, K outlet gas temperature, K outlet gas temperature, K refractory or insulating temperature for layer ‘n’, K reference temperature, K internal wall temperature: refractory surface, K wall temperature of the spiral exchanger, K wall temperature at section ‘n’ of the spiral exchanger, K total internal energy, J outlet gas velocity, m/s combustion chamber volume, m3 outlet air flowrate, kmol/s inlet air flowrate, kg/s auxiliary fuel flowrate, kg/s inlet gas flowrate, kmol/s outlet gas flowrate, kmol/s inlet component ‘i’ flowrate deriving from inlet flows or solid/gas reactions, kmol/s outlet gas flowrate, kg/s outlet solid flowrate containing ash, inert or unburned products, kg/s inlet waste flowrate, kg/s oxygen bulk mole fraction in the gas phase oxygen interface mole fraction assumed equal to zero over all the flame surface Greek letters a waste nitrogen fraction converted to NO according to Bowman (1975) b dynamic angle of repose for the solid material c geometry coefficients i 2742 e p e G h k O2 l O2 o o W p q u BF u H2O u O2 u N2 M. Rovaglio et al. refractory emissivity gas emissivity residence time, s stoichiometric coefficient for waste combuswhere the waste refers to tion, kg /kg O2 8!45% the combustible fraction and the O con2 sumption refers to CO formation stoichiometric coefficient for waste combuswhere the waste refers tion, kmol /kmol O2 8!45% to the combustible fraction and the O con2 sumption refers to CO formation gas density, kg/m3 wall conductivity, W/m K Stefan—Boltzmann constant, W/m2 K4 gas transmittance burning mass fraction in the inlet waste water mass fraction in the inlet waste (moisture) oxygen mass fraction in the inlet air flowrate nitrogen mass fraction in the inlet air flowrate REFERENCES Behmanesh, N., Manausioithakis, V. and Allen, D. T. (1990) Optimizing the throughput of hazardous waste incinerators. A.I.Ch.E. J. 36, 1707—1714. Biardi, G., Pellegrini, L., Grottoli, M. G. and Rovaglio, M. (1987) Chemical furnaces under transient conditions. Proceedings of CEF ’87, Taormina, Italy. 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