4708
Ind. Eng. Chem. Res. 2003, 42, 4708-4714
KINETICS, CATALYSIS, AND REACTION ENGINEERING Experiments and Simulations of Time-Dependant Phenomena in a Hydrothermal Oxidation Tubular Reactor Patrick Dutournie´ Laboratoire d’Etude Thermique, Energe´ tiques et Environnement, Universite´ de Bretagne Sud, Centre de recherche, 56100 Lorient, France
Cyril Aymonier and Franc¸ ois Cansell Institut de Chimie et de la Matie` re Condense´ e de Bordeaux (ICMCB), Universite´ Bordeaux I, 87 avenue du docteur Schweitzer, 33608 Pessac Cedex, France
Jacques Mercadier* Laboratoire de Thermique, Energe´ tique et Proce´ de´ s (LATEP), Ecole Nationale Supe´ rieure en Ge´ nie des Technologies Industrielles, rue J. Ferry, 64000 Pau, France
The scaling up of a hydrothermal oxidation (HTO) facility depends on the development of simulation tools. In this way, a new approach to simulate the reactor thermal behavior during a nonstationary regime such as the start, stop, or emergency stop of the unit is developed. A numerical procedure of nonstationary temperature profile simulation in the reactor is described in order to predict the reactor thermal behavior. This new model is validated from experimental data obtained with a quasi-adiabatic reactor, which allows one to follow the temperature profile evolution along the reactor during the HTO process. Performances of the numerical procedure are tested with the simulation of three kinds of real nonstationary conditions: facility start, pump defect, and waste concentration dilution. The simulation gives the evolution of temperature profiles of a reactive mixture versus time and models the external wall temperature of the reactor tube versus time. These three simulations show that the HTO reactor thermal operation can be considered as stable and sure. Furthermore, they put in a prominent position that the risks of reactor damage are limited. I. Introduction Hydrothermal oxidation (HTO) concerns oxidation of organic matter in water under high pressure and high temperature (P > 22.1 MPa and T > 374 °C). The principle and performances of HTO processes are today well-known and have already been described elsewhere.1 This process is being developed as a technique for treating liquid wastes and sludge and allows one to achieve these major aims: energetic utilization, degradation in environmentally acceptable end products, and final volume reduction of wastes and sludges. The first commercial plant facility of HTO has been operating since 1994 in Austin, TX, treating 1000 L h-1 of liquid effluent (about 4%w waste). Its efficiency is greater than 99.5%. Since 1994, other HTO facilities have been built, especially in Japan and USA. Some of them are used for sludge treatment (Hydroprocessing, 2001, Harlingen, TX). Today the development of HTO processes depends principally on the conception of new reactor concepts2 and on the development of simulation tools, which are necessary for scaling up HTO processes.
Specific reactors have already been studied from trading simulation computer softwares3-5 (such as FLUENT or the Computational Fluid Dynamics model), for example, in nonturbulent flows. Other works6-8 were devoted to the model of stationary or nonstationary concentric-tube reactors in turbulent flow. To avoid irreversible reactor damage and to ensure the security of the process, it seems very important to control the temperature, especially during transition phases (unit start, emergency stop, etc.). However, most of simulation tools operate in steady state, and models in nonstationary phenomena are limited. From this point of view, a model that can simulate the reactor’s thermal behavior in the non steady state is presented. This model is validated from data obtained with a continuous pilot-plant facility. After the description of the pilot, experimental results, and numerical program, the last part of this paper consists of a comparison between experimental and theoretical temperature profiles of different nonstationary phenomena. II. Experimental Section
* To whom correspondence should be addressed. Fax: 33.5.59.40.78.01. E-mail:
[email protected].
Quasi-adiabatic Tubular Reactor. The quasiadiabatic tubular reactor (Figure 1) is made of an
10.1021/ie0300732 CCC: $25.00 © 2003 American Chemical Society Published on Web 09/04/2003
Ind. Eng. Chem. Res., Vol. 42, No. 20, 2003 4709
Figure 2. Steady-state temperature profiles along the reactor.
Figure 1. Quasi-adiabatic tubular reactor.
Table 1. Pilot Facility Operation Conditions and Oxidation Results test waste 1 2 3 4 5
1 2 3 4 4
m ˘w m ˘ ox outlet COD Tmax inlet COD (kg m-3) (kg h-1) (kg h-1) (10-3 kg m-3) (°C) 57.8 63 81.5 78.7 78.7
0.5 0.55 0.5 0.55 0.55
0.5 0.45 0.4 0.55 0.45
20 28 4 40 5
572 565 563 572 565
Inconel 718 tube of 2.4 mm inside diameter and 8.4 m length (volume of 38 mL). It is equipped with 28 thermocouples (type K) distributed along the reactor. The tube is insulated with ceramic fiber and surrounded with a thermal shield. The thermal shield consists of a copper sheet surrounded with an electrical heater in order to reduce thermal loss. The quasi-adiabatic tubular reactor allows one to obtain a temperature profile along the tube that characterizes the reactor thermal behavior. At the reactor outlet and after pressure reduction, the composition of the gas phase is analyzed by inline gasphase chromatography. The organic amount in both waste feed and liquid effluent is determined by measurement of the chemical oxygen demand (COD) with a Behrotest TRS 200 using a colorimetric method.9 III. Thermal Data Wastewaters were injected in the reactor at 400 °C and 25 MPa, and the oxidant, hydrogen peroxide, was injected in the reactor at room temperature in order to avoid decomposition during preheating.10 The hydrogen peroxide concentration of the oxidant feed solution was 166 g L-1, and the flow rate was adjusted from the waste flow rate and COD in order to bring H2O2 in the reactor into stoichiometry (Table 1). The residence time of the reactive mixture (wastewater + oxidant solution) was calculated to be about 30 s. Labview was used to control facility operations. The reactor’s thermal behavior was studied from four real industrial wastewaters (wastes 1-4). The exact compositions of these wastewaters are not well-known;
they are composed of organic compounds (C, H, and O), and their initial CODs are presented in Table 1. The operating conditions of the pilot facility (oxidant and waste mass flow rates [m ˘ ox and m ˘ w]), the COD at the facility outlet, and the maximal temperature (Tmax) reached in the reactor are summarized in Table 1. Figure 2 shows the different temperature profiles (tests 1-5) obtained with the quasi-adiabatic reactor working in steady state. The evolution of the temperature profiles versus the reactor length can be divided into two parts. The first one (reactor length < 2 m) corresponds to the exothermic degradation of the organic matter of the waste, and in the second part (reactor length > 2 m), the oxidation reaction is quasi-complete and the heat released by the reaction quasi-null; this second part of the reactor allows one to optimize the residence time to complete destruction of the organic matter. During the different tests, the H2O2 concentration and flow rate are constant and the stoichiometry is conserved by adjusting the waste flow rate. Hence, the carbon concentration in the global mass flow rate is approximately constant, leading to a quasi-similar temperature profile. Furthermore, the second part of the temperature profiles shows that the heat loss along the reactor can be neglected. During the oxidation run of test 1, the reactor’s thermal behavior was investigated in starting the reaction. The evolution of the temperature profiles along the reactor between 30 s after the waste feed injection in the reactor ()start) and 15 min after the facility start is shown in Figure 3. Before the waste injection, pure water was injected at 400 °C in the waste injection line and H2O2 is injected at room temperature in the oxidant injection line with a mass flow rate corresponding to the one of the experiment (the one of test 1). When pure water is replaced by the waste, the oxidation reaction takes place and the temperature increased rapidly in the first 2 m of the reactor as explained previously. In these experiments, an increase of about 150 °C was observed and steady state was reached after 15 min of operation. A total of 8 min after a facility stop (see also Figure 3), the temperature profile is quasi-similar to steady state at the reactor inlet (between 0 and 0.75 m) and at the reactor outlet (between 3 m and the reactor end). A difference of approximately 100 °C can be observed
4710 Ind. Eng. Chem. Res., Vol. 42, No. 20, 2003
∂ (Fu) ) 0 ∂z
(1)
u(0) ) m ˘ /SF(T0) ∂u ∂u ∂P 4 ∂ µ )+ ∂z ∂z 3 ∂z eff ∂z
(
Fu
)
(2)
P(0) ) P0 ∂uCOD ∂2COD -k ) Deff ∂z ∂2z
(3)
C(0) ) C0 Figure 3. Evolution of the temperature profile for the steady state (*), 15 min after start of the reaction (O), and 8 min after stop of the reaction (4). [: reactor start.
FuCp
∂T uT ∂F ∂P ∂ ∂T + ) λ - R∆Hk + q˘ (z) ∂z F ∂T P ∂z ∂z eff ∂z (4)
( )
(
)
T(0) ) T0 q˘ (z) )
|
2λm ∂Tm (thermal flux equality R1 ∂r r)R1 between the fluid and metal material)
Energy Conservation in the Metal (r ) R1 to R2).
( )
∂Tm ∂2Tm 1 ∂ r + )0 r ∂r ∂r ∂z2
Figure 4. Schematic description of the reactor computational domain.
Tm(R1,z) ) T(z) Tm(r,0) ) T0
between a reactor length of 0.75 and 3 m, corresponding to the reactor part (between 300 and 550 °C) where most of the oxidation heat is released, making this part of the reactor very sensitive to the oxidation reaction. As expected, a facility stop, such as, e.g., a feed pump defect, would not disturb the process stability and security. These tests will be used to validate the model. IV. Numerical Procedure The numerical procedure was focused on the tubular reactor, in which the flow regime is generally turbulent. This hydrodynamic regime has for principal effect to increase exchanges in the fluid.11 A previous study12 has shown that this kind of reactor is insensitive to body force effects. As expected, the simulated12 profiles have shown that there is no radial temperature and concentration gradient and that the velocity profile is flat. This consideration allows the use of a one-dimensional model to study these kinds of reactors. Moreover, this study has shown that the contribution of the turbulence in the equation of energy can be neglected by a report of the heat produced by the chemical reaction. Thus, to simplify the program, no turbulence model was used. In fact, the effects of the effective viscosity, thermal conductivity, and turbulence heat are insignificant. IV.1. Simulation in Steady State. The numerical model requires a simultaneous resolution of momentum, mass balance, energy, and chemical species conservation equations in the fluid and of the energy conservation equations in the metal and in the insulated material. Flow Equations (z ) 0-L). A schematic description of the configuration of the reactor is given in Figure 4.
(5)
∂Tm ∂z λm
|
z)L
|
∂Tm ∂r
r)R2
)0
) λi
|
∂Ti ∂r
r)R2
Energy Conservation in the Insulated Material (r ) R2 to R3).
( )
∂2Ti 1 ∂ ∂Ti r + 2 )0 r ∂r ∂r ∂z
(6)
Ti(r,0) ) T0 Ti(R3,z) ) Tp
|
∂Ti ∂z λm
|
∂Tm ∂r
z)L
)0
|
∂Ti (heat flux equality ∂r r)R2 between the metal and insulated material)
r)R2
) λi
This part of the program allows one to determine the steady-state thermal behavior of the reactor. The different stationary profiles will be used as initial conditions in the nonstationary simulations. IV.2. Simulation in the Nonstationary Regime. The equation system is the same as the one of the stationary system with unsteady contributions.
Ind. Eng. Chem. Res., Vol. 42, No. 20, 2003 4711
Flow Equations (z ) 0-L).
Material properties
∂F ∂ + (Fu) ) 0 ∂t ∂z
(7)
FCp
0.01 W m-1 K-1 (200-900 °C)
∂u ∂u ∂u ∂P 4 ∂ F + Fu )+ µeff ∂t ∂z ∂z 3 ∂z ∂z
)
(8)
∂2COD ∂COD ∂uCOD -k + ) Deff ∂t ∂z ∂2 z
(9)
(
V. Simulation and Model Validation in the Steady State
∂T ∂T T ∂F ∂P ∂P ) + FuCp + +u ∂t ∂z F ∂T P ∂t ∂z ∂ ∂T 2 λ - R∆Hk + q˘ (z) (10) ∂z eff ∂z
( )[ ( )
]
Energy Conservation in the Metal (r ) R1 to R2).
FmCpm
[
]
( )
∂Tm ∂2Tm ∂Tm 1 ∂ r + ) λm ∂t r ∂r ∂r ∂z2
(11)
Energy Conservation in the Insulated Material (r ) R2 to R3).
FiCpi
[
( )
]
∂Ti ∂2Ti ∂Ti 1 ∂ r + 2 ) λi ∂t r ∂r ∂r ∂z
(12)
Solving the Differential Equation System. This system of six equations is solved simultaneously by using the same boundary conditions as those defined in the steady-state simulation. This code permits one to modulate the inlet COD (COD0), the inlet temperature (T0), or the inlet mass flow rate (m ˘ ). The velocity, temperature, and concentration profiles are modeled by using an integral method (Euler) with a constant time step and the spatial variations by using the finite difference method. IV.3. Simulation of Chemical Reactions. The simulation of chemical reactions was developed for steady and unsteady states. According to the literature,13,14 first-order kinetics is used:
k ) RAe-Ea/RTCOD
Deff ) 10-6 m2 s-1, λm ) 15 W m-1 K-1, λi ) 0.00027(T - 273.15) -
The simulations were run at the same conditions as those of test 1 for the facility start at t0 (around 37 s) and after 15 min. At t0, simulated and experimental temperature profiles of the reactor tube are in good agreement (Figure 5). However, the simulation shows that the temperature of the reactive mixture increases very rapidly at the injection point of waste and oxidant feed. This temperature is different from that of the reactor tube because of the thermal inertia of the system. This difference decreases rapidly to become negligible after 15 min. The simulated temperature profile is then close to the experimental one and justifies the validity of the numerical procedure. Figure 5 shows that the conclusion is the same for the established profiles 8 min after a pumps stop (waste and oxidant). Indeed, the characteristic time of heat transfer in the reactor tube alloy is lower than 1 min (the conduction time in the material is around t ) L2FCp/λ ≈ 15 s). The average residence time has the same order of magnitude (less than 30 s for a reactor length < 5 m), but the conduction time in the ceramic fibers is higher than 1
(13)
A numerical program15 had been developed to calculate the reaction heat and kinetic parameter by use of experimental results and numerical simulations. These are
R∆H ) 40.27 MJ kg-1, A ) 2970 s-1, Ea ) 51.1 kJ mol-1 IV.4. Operating Conditions. In the numerical procedure, the fluid is considered to be pure water from a thermodynamic point of view. “The 1967 IFC Formulation for Industrial Use”16 was used to determine the density, enthalpy, and specific heat of pure water. Thermal conductivity and dynamic viscosity were calculated with specific equations.16
Operating conditions P0 ) 25 MPa, T0 ) Tp ) 200 °C, ˘ ) 1 kg h-1 DCO0 ) 57.8 kg m-3, m
Figure 5. Experimental and simulated temperature profiles (waste and reactor wall) along the reactor at t0 (waste injection; a), 15 min after the waste injection (b), and 8 min after the pumps stop (c).
4712 Ind. Eng. Chem. Res., Vol. 42, No. 20, 2003
Figure 7. Temperature profiles of reactive mixture (thick line) and wall reactor (thin line) versus time after pumps stop. Figure 6. Temperature profiles of fluid (thick line) and wall reactor (thin line) versus time at facility start.
h. So, an important difference of temperature can occur between the reactive fluid mixture and the external wall of the reactor tube, as is shown in the first part of Figure 5, especially during the first seconds after a non-steadystate modification of the facility operation. VI. Simulation in a Nonstationary Regime After model validation, three kinds of unsteady conditions are studied in this section: facility start, pumps stop, and dilution of the waste concentration. VI.1. Reactor Start. For the simulation of the facility start, the temperature of the fluid in the reactor is assumed to be equal to 200 °C and the oxidant’s injection corresponds to t ) 0.
Figure 6 presents the temperature profile evolution of the fluid and the reactor wall during 50 s after the oxidant injection. Thermal equilibrium is reached after 2 min. Before the steady state is reached, a thermal gradient of up to 200 °C can evolve between the reactive fluid and external wall of the reactor tube. During the first 2 min, a maximum temperature is reached, which may lead to thermal strain. VI.2. Reactor Stop. Figure 7 presents the temperature profile evolution of the fluid and the reactor wall after a pump stop, assigned as t ) 0, being the steady state. Thermal equilibrium is obtained only after several hours because heat losses across the ceramic fibers are low. When the pumps are stopped, the reactor becomes a closed system, the flow rate is 0, and heat production cannot be dissipated (except by volumetric expansion). Conversion in the steady state is complete, and the
Ind. Eng. Chem. Res., Vol. 42, No. 20, 2003 4713
more important than the one produced in the steady state at the middle of the reactor. VII. Conclusion A numerical procedure of simulation has been developed by LaTEP. The simulation of a non-steady-state regime of a HTO reactor allows one to predict the temperature profiles of the fluid and the reactor material along the reactor. The model used has been validated from experimental temperature profiles obtained with a quasi-adiabatic HTO tubular reactor. Temperature profiles of four different real industrial wastes have been measured, and the evolution of the temperature profiles of one of them was followed during the facility start and pump stop. A good agreement between the experimental and simulated temperature profiles along the reactor was obtained. The performances of the validated numerical procedure have been tested with the simulation of three kinds of real nonstationary conditions: facility start, pump defect, and waste concentration decrease. The simulation gives the evolution of the temperature profiles of the fluid and the reactor wall versus time. These three simulations show that the HTO reactor thermal operation can be considered to be stable and sure. They put in prominent position that the risks of reactor damage are limited and that the developed model is able to prevent the problems in regards to the reactor’s thermal behavior. This numerical procedure of simulation in the non-steady-state regime of HTO reactor thermal behavior is a tool necessary for scaling up HTO facilities. Appendix
Figure 8. Temperature profiles of fluid (thick line) and wall reactor (thin line) versus time after a decrease of the waste concentration.
temperature increase at the beginning of the reactor is inferior to the maximum temperature reached (480 °C compared to 570 °C). VI.3. Waste Dilution. Figure 8 presents the temperature profile evolution of the fluid and the reactor wall after a decrease of the waste concentration. The reactor works in the steady state, and at t ) 0, pure water is injected into the reactor in order to decrease the waste concentration by a factor 1.7 at the reactor inlet. The simulation shows that the fluid temperature decreases faster than the one of the reactor wall. However, an increase of the fluid temperature after 5 s is observed. This phenomenon can be explained by a decrease of the conversion in the first reactor part. In fact, the temperature decrease at the reactor inlet induces a decrease of the oxidation reaction rate. The heat that had remained in the reactor wall becomes
A ) preexponential constant (s-1) COD ) chemical oxygen demand (mg L-1) Cp ) specific heat (J kg-1 K-1) D ) effective diffusion coefficient of a compound in the mixture (m2 s-1) Ea ) activation energy (kJ mol-1) k ) reaction rate (mol m-3 s-1) L ) length of the reactor (m) m ˘ ) inlet mass flow rate (kg s-1) P ) static pressure (Pa) q˘ ) exchanged heat flux at the wall by unity of volume (W m-3) r ) radius of the reactor (m) R1 ) internal radius of the annular material R2 ) external radius of the annular material R3 ) external radius of the insulated material S ) reactor section (m2) T ) temperature (K) u ) velocity (m s-1) z ) coordinate of the reactor axis (m) Subscripts i ) initial or inlet f ) final or outlet 0 ) inlet m ) material w ) waste p ) thermal shield ox ) H2O2 feed solution Greek Symbols R ) ratio (R ) C/COD) ∆H ) reaction enthalpy (J kg-1) λ ) thermal conductivity (W m-1 K-1)
4714 Ind. Eng. Chem. Res., Vol. 42, No. 20, 2003 µeff ) effective dynamic viscosity (kg s-1 m-1) F ) density (kg m-3)
Literature Cited (1) Savage, P. E.; Gopalan, S.; Mizian, T. I.; Martino, C. J.; Brock, E. E. AIChE J. 1995, 41, 1723. (2) Aymonier, C.; Bottreau, M.; Berdeu, B.; Cansell, F. Ultrasound for hydrothermal treatments of aqueous wastes: solution for overcoming salt precipitation and corrosion. Ind. Eng. Chem. Res. 2000, 39, 4734. (3) Oh, C. H.; Kochan, R. J.; Beller, J. M. Numerical analysis and data comparison of a supercritical water oxidation reactor. Environ. Energy Eng. 1997, 43 (6), 1627. (4) Oh, C. H.; Kochan, R. J.; Charlton, T. R. Modelling of thermal characteristics in supercritical water oxidation reactors. Proc. ASME Heat Transfer Div. 1995, HTD-317-2, 311. (5) Michelfelder, B.; Noll, B.; Weindel, M.; Eckl, W.; Eisenreich, N.; Herrmann, M. M. A high-pressure combustion cell based on numerical flow simulation and reaction zone radiation modelling. In High-Pressure Chemical Engineering; Rudolf von Rohr, Ph., Trepp, C., Eds.; Elvesier Science BV: Amsterdam, The Netherlands, 1996. (6) Chen, P.; Li, L.; Gloyna, E. F. Simulation of a concentrictube reactor for supercritical water oxidation. Innovations Supercrit. Fluids 1995, 24, 348. (7) Petrich, G.; Abeln, J.; Schmieder, H. Model and simulation of supercritical water oxidation. High Pressure Chem. Eng. 1996, 157. (8) Dutournie´, P.; Mercadier, J. One-dimensional simulation of a tubular reactor for waste oxidation. Global Symposium on
Recycling, Waste Treatment and Clean Technology; INASMET Publisher: San Sebastian, 1999; p 1875. (9) Aymonier, C.; Beslin, P.; Jolivalt, C.; Cansell, F. Hydrothermal oxidation of a nitrogen-containing compound: the fenuron. J. Supercrit. Fluids 2000, 17, 45. (10) Beslin, P. Conversion hydrothermale des de´chetssEtude de la re´activite´ chimique de compose´s mode`les et applications a` la detruction de boues d’effluents indutriels. Ph.D. Thesis, University of Bordeaux, Pessac Cedex, France, 1997. (11) Cansell, F.; Delville, M. H.; Subra, P. Fluides Supercritiques et Mate´ riaux; ISASF Publisher: Nancy, France, 1999. (12) Dutournie´, P.; Mercadier, J.; Ce´zac, P. Simulation of a tubular reactor for supercritical water oxidation. Re´ cent Prog. Ge´ n. Proc. 1999, 8, 70, 407. (13) Li, L.; Chen, P.; Gloyna, E. F. Kinetic model for wet oxidation of organic compounds in subcritical and supercritical water. Supercrit. Fluid Eng. Sci. 1993, 24, 306. (14) Webley, P. A.; Holgate, H. R.; Stevenson, D. M.; Tester, J. W. Oxidation Kinetics of model compounds of metabolic waste in supercritical water. SAE Technical Paper Series 901333; SAE: Williamsburg, VA, 1990. (15) Dutournie´, P.; Mercadier, J.; Aymonier, C.; Gratias, A.; Cansell, F. Determination of hydrothermal oxidation reaction heats by experimental and simulation investigations. Ind. Eng. Chem. Res. 2001, 40, 1, 114. (16) Schmidt, E. Properties of Water and Steam in SI-Units; Springer-Verlag: New York, 1989.
Received for review January 27, 2003 Revised manuscript received July 18, 2003 Accepted July 22, 2003 IE0300732
