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Development of a non-equilibrium turbulent reaction model of HFC-134a in a tubular reactor
Applied Thermal Engineering 125 (2017) 20–28
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Research Paper
Development of a non-equilibrium turbulent reaction model of HFC134a in a tubular reactor Misoo Shin a, Dongsoon Jang a,⇑, Jongwook Ha b, S. Acharya c a
Department of Environmental Engineering, Chungnam National University, 99 Daehak-ro, Yuseong-gu, Daejeon 34134, South Korea Korea Research Institute of Chemical Technology, 141 Gajeong-ro, Yusunggu, Daejeon 305-600, South Korea c Department of Mechanical Engineering, Louisiana State University, Baton Rouge, LA 70803, USA b
h i g h l i g h t s A hybrid-type EBU model is developed for HFC-134a incineration. Experiment and numerical prediction data are well compared each other. The hybrid EBU model shows big difference with the conventional model.
a r t i c l e
i n f o
Article history: Received 18 October 2016 Revised 23 May 2017 Accepted 20 June 2017 Available online 22 June 2017 Keywords: Halogen compound incineration HFC-134a Non-equilibrium turbulent reaction model Hybrid-type eddy breakup model
a b s t r a c t Detailed review and discussions are made for the modeling and non-equilibrium effect in the incineration process of various halogenated hydrocarbons using a couple of modified versions of eddy breakup (EBU) model developed from the original EBU model by Magnussen and Hjertager. Further experimental and numerical studies are made for the development of a hybrid-type EBU model in order to evaluate the non-equilibrium effect of HFC-134a reaction in a tubular reactor. For the numerical calculation using the hybrid-type EBU model, especially, the relative rate of the turbulent mixing and chemical reaction should be incorporated in a harmonic mean expression. Thus the overall reaction rate expression of HFC-134a reaction is determined by the fitting the empirical constants with experimental data obtained in the tubular reactor via trial and error method. Thereby, the empirical constants in the equation of the E _ ¼ kC C2 H2 F4 C H2 O C O2 ¼ AT B eRTa C C2 H2 F4 C H2 O C O2 was determined, Arrehenius type reaction rate, that is, x where A = 1.0, B = 0.38, Ea = 8.8 107 J/kmol, respectively. Employing this non-equilibrium model, the calculated overall destruction rate of HFC-134a was in good agreement with the experimental data obtained for the same reactor. Further, the comparison of HFC-134a concentration profiles between the fast chemistry and non-equilibrium EBU models show a significant difference even if the calculated destruction rates of HFC-134a appear similar at the exit. In near future, numerical calculation will be made for the practical incinerator of HFC-134a in a practical incinerator using the model developed in this study. Ó 2017 Elsevier Ltd. All rights reserved.
1. Introduction These days, reclaim and recovery services of used refrigerants such as HFCs are one of major concern in order to minimize the environmental impact of waste refrigerants especially in the aspect of the release of the material of high global warming potential. However, a significant part of recovered refrigerants have to be destroyed inevitably in a practical incinerator in an environmentally friendly manner when their compositions or purity are inferior than specified condition for reclamation [1]. As a matter ⇑ Corresponding author. E-mail address: [email protected] (D. Jang). http://dx.doi.org/10.1016/j.applthermaleng.2017.06.105 1359-4311/Ó 2017 Elsevier Ltd. All rights reserved.
of fact, the incineration has been known over several decades as a preferred method for hazardous waste disposal compared to other conventional methods because of its potential for immediate detoxification and the potential for energy recovery. A proper incineration method for the waste HFCs (hydroflurocarbons) has become a viable method not only for the reduction of the release of the greenhouse gases but also for the recovery of the valuable material such as CaF2 (calcium fluoride) [2]. However, it is generally known the practical incineration process of most refrigerants containing halogen species such as F and Cl could not be properly modeled by typical fast chemistry turbulent reaction model due to not only its flame inhibition effect but also the low heating value of halogen containing species. Thus a proper non-equilibrium
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M. Shin et al. / Applied Thermal Engineering 125 (2017) 20–28
turbulent reaction model is quite required for a practical incinerator to develop a reliable predictive model in order to obtain proper design and operational conditions for the practical HFCs incineration in turbulent reacting flows. To this end, in this study detailed review and discussions are made first for the modeling and nonequilibrium effect in the incineration process of various halogenated hydrocarbons. And a numerical and experimental study has been made in a tubular reactor for the development of proper non-equilibrium turbulent reaction model of HFC-134a.
2. Review of halogenated compound incineration and numerical modeling Research on the combustion of various halogenated hydrocarbons has been done mainly in incineration area of hazardous waste, for chlorinated hydrocarbons, generated from chemical industries since 1970. Since incineration is known as a preferred tool for hazardous waste disposal compared to other conventional methods, this technology is also certified by United Nations Framework Convention on Climate Change (UNFCC) for the thermal destruction of halogenated hydrocarbons. However, as well known, only general guideline is given to accomplish 99.99% thermal destruction efficiency (DRE), based on the empiricism, suggesting that it requires flame temperature higher than 1200 °C and the residence time more than 2 s [3]. Since halogenated hydrocarbons not only serve as fire extinguishing agents, but also some of those cannot produce self-sustaining flame. One of the typical ways for proper incineration, therefore, is to co-burning of this compound with auxiliary fuels of containing many hydrogen atoms such as CH4 and H2. Halogenated hydrocarbons with the aid of this kind of fuel can be effectively decomposed by the reaction with hydrogen and oxygen; hydrogen combines with halogen to form HCl or HF. In this process, it is noted that the halogen atom originated from halogenated hydrocarbon plays not only the role of oxidizer but also inhibition action through the combination of the hydrogen containing in auxiliary fuel. In this process, the conventional strategy of introducing excess air to ensure complete combustion of the waste stream in practical incinerators is not quite desirable to the incineration of CHCs (Chlorinated Hydrocarbons) because of the following Deacon reaction [4].
1 2HCl þ O2 ! Cl2 þ H2 O 2
ð1Þ
Excess oxygen can lead to pure chlorine gas, Cl2, which is more toxic and difficult to scrub out of stack gases than HCl [5]. This reaction together with the inhibition effect makes incineration of CHCs difficult to accomplish the goal of DRE 99.99%. Similar argument can be applied to the HFCs refrigerants containing fluorine. In general, the fundamental studies have shown that the incineration process of halogenated hydrocarbons is usually inhibited by the reaction of halogen molecule (X2) with hydrogen H especially in preheating region since the activation energy of inhibition reaction is much lower than that of the chain branching reaction [6]. Garner et al. [7] and Valeiras [8] showed that the order of effectiveness of in inhibition for chloromethane by decreasing flame velocity is CCl4, CHCl3, CH2Cl2 and CH3Cl with the increase of the chlorine molecule. The flame velocities of the mixture are, of course, lower than methane and decrease with the increase of the molar fraction of chlorine compound to the auxiliary fuel (R), for example R = mole of CHCs/mole of CH4 as in Eq. (2).
CH4 þ RCCl4 þ 2O2 þ 7:25N2 ! ð1 þ RÞCO2 þ 2ð1 RÞH2 O þ 4RHCl þ 7:25N2
ð2Þ
Further as the chlorine content increases by R or by the change compound, the flammability range decreases and the maximum flame velocity shifts from fuel rich to fuel lean condition [9]. The net effect of the halogen inhibition on flame is to increase the length of the preheat zone over that of the uninhibited flame but the higher temperature in the reaction zone increases exponentially the chain branching reaction rate enough to overcome the rates of the inhibition reaction [7]. Further, iodine and bromine compounds are found to be more strong inhibitors than chlorine and fluorine compounds [6]. Even if HFCs have been developed as essential substitute for the ozone depleting refrigerants or sprays such as CFCs and HCFCs containing chlorine compounds, some of the alternative HFCs show unexpectedly high flammability. Therefore, in order to ensure fire safety application of these substances, there has been a lot of fundamental combustion study about fluorinated hydrocarbons [8–14]. A detailed reviews about this have been made elsewhere [15]. A series of detailed chemistry data of HFC-134a appear only in the research area of degradation pathways into trifluoroacetic acid (TFA) in the atmosphere instead of combustion field like waste incineration. In detail, the OH radical is expected to be the most important oxidation of HFCs degradation in daytime troposphere. Extensive experimental studies have been taken on the kinetics of the hydrogen abstraction of HFC134a by hydroxyl radical since this reaction plays an important role in the gas-phase reaction of HFC-134a in the atmosphere [16]. The rate constants determined are presented but unfortunately the temperature ranges are below than 500 K, which is far below the flame temperature. Along with the fundamental combustion study, even though a lot of research efforts [17–19] have been made about thermal destruction of halogenated compound both in large scale incinerators and also model developments have been made to predict the performance of incineration processes, these studies do not provide a predictive tool for resolution of the detailed distributions of the turbulent flow field. Even though Yang et al. [20] made a detailed numerical calculation for a practical rotary kiln incinerator, they did not include any halogenated compounds as principal organic hazardous constituents. In general, there are two fundamental steps for the process of turbulent combustion; the first step is turbulent mixing between fuel and air, and the next chemical reaction by the molecular collision between fuel and oxidizer. In order to make complete combustion successfully, therefore, it is necessary to mix efficiently fuel with air by eddy breakup action and then fuel molecule will react with oxygen molecule for chemical reaction. The overall reaction rate occurring in a series of competitive process can be typically expressed empirically in a harmonic form such as
Ov erall reaction rate
1 1 turbulent mixing rate
1 þ chemistry
ð3Þ rate
The basic problem of the gaseous turbulent combustion lies in the proper modeling of the well-known turbulent correlation term, which arises in the non-linear chemical reaction source term like Reynolds stress in momentum equation. Most turbulent combustion models, however, detour the difficulty by ignoring the chemical kinetic aspects of the problem as shown in the harmonic mean expression of overall reaction rate in Eq. (3). In detail, the chemistry rate is fast and thereby very large compared to turbulent mixing rate. Therefore the term can be neglected. Using the assumption of fast chemistry, the turbulent mixing will be governed only by the limit of the mixing rate in Eq. (3). Although mixing limited turbulence only models have been shown to be successful tools in predicting the overall profiles of main species, serious drawback exist, since the conventional fast chemistry turbulent reaction models such as the conserved scalar method of Bil-
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ger [21], the eddy breakup models of Spalding [22] and Magnussen and Hjertager [23] do not describe the non-equilibrium characteristics of halogen compounds. Since the effect of inhibitive chemistry reaction is as important as that of turbulent mixing an appropriate turbulent combustion model had to be used, where chemical kinetics as well as turbulence mixing plays an important role. In the past, considerable progress has been made in the modeling of chemical kinetic processes in turbulent flows for various areas such as flame extinction in strong shear flows, NOx generation in pollutant formation and chlorination process in chemical reactor [24–28]. However, general non-equilibrium turbulent reaction model is not yet available especially for the halogenated incineration processes. For the numerical modeling efforts of CHCs, CCl4 destruction has been done mainly since CCl4 was one of the major hazardous waste to be incinerated. Therefore a number of investigators have focused on the modeling of the turbulent reaction in a premixed CCl4-CH4-air flame. As emphasized above, the reaction rate of this mixture is believed to be determined not only by turbulence mixing but also by chemical kinetics. This is particularly, true since CCl4 is incapable of a self-sustaining flame due to inhibition feature as well as the low enthalpy of combustion. In order to resolve the non-equilibrium effects in CCl4-CH4-air flames, detailed chemical kinetic data is required. And a number of research efforts have been directed on this end [17], but there was no wellestablished chemical kinetic data available. Thus an empirical modeling approach was suggested in the CCl4 incineration process. The first reduced reaction model for CCl4 destruction [29] is presented briefly in the following for the discussion of the modified version of EBU model. The reaction rate of CH4 (RRCH4 ) in a CCl4CH4-air mixture is given by the phenomenological eddy breakup model proposed by Magnussen and Hjertager [23], i.e.
this turbulent reaction model. Shin et al. [31] employs a similar eddy breakup model for the reaction of the mixture of CH4-CCl4. But they consider the mixture of CH4-CH3Cl behaves together as fast chemistry reaction, while Jang and Acharya [29] assume that only CH4 reaction is governed by fast chemistry. The reason why Shin et al. employ the fast chemistry model for the total mixture of CH4-CCl4 was in that the small amount of CCl4 would not play ant significant inhibition action most over strong turbulent reaction region. This model is reasonable only for the case of the small amount of CCl4 presence. Further, Jang and Acharya [32] propose a non-equilibrium turbulent reaction model using burning velocity data given by Valeiras [8] for CH4-CCl4-air mixture e in 2-D hypothetical kiln. That is, a modified fast chemistry turbulent reaction model was developed to describe the flame inhibition due to the presence of CCl4, considering the corresponding burning velocity data of these mixtures. This modified EBU model was employed with reasonably good prediction for numerical calculation of practical incinerators for a rotary kiln and dump combustor [33–35]. The experimental burning velocity data of premixed CCl4-CH4-air mixture as a function of R and equivalence ratio in a laboratory Bunsen flame study [6] is utilized to obtain a more realistic reaction rate expression. In the aspect to review the modified eddy breakup model for non-equilibrium turbulent combustion of halogenated hydrocarbons, let us briefly summarize this model. Based on the thermal theory of flame propagation [36], the reaction rate (RR) can be related to the flame burning velocity (Su ) by
pr m m Am fu =k; q A0 OX ð=kÞ; qA0 RRCH4 ¼ minimum of q s ð1 þ sÞ k
2 Su a R; / RRCH4 R¼0;/ RRCH4CCl4 at R;/ ¼ 2 Su a R ¼ 0; /
ð4Þ where s is the stoichiometric oxidizer mass per unit mass of fuel and A0 and A are empirical constants given by Lockwood et al. [30]. The basic idea of this model is based on premise that only the reaction of CH4 in a CCl4-CH4-air mixture is governed by the fast chemistry concept and the destruction of CCl4 can be described by the stoichiometric reaction Eq. (2) according to the destruction of CH4. It is generally expected that the CH4 reaction rate given by Eq. (4) will be smaller compared to that in a pure CH4-airmixture, (i.e. R = 0.0) because of the decreased mass fractions due the presence of CCl4 and HCl compound. In order to account for the CH4 reaction rate in CCl4-CH4-air mixture, the following expressions arc used for the mass fractions appearing in Eq. (4)
mfu ¼ mCH4 ; mOX ¼ mO2
and mpr ¼ mCO2 þ mH2 O þ mHCl
ð5Þ
Once the CH4 reaction rate is obtained, the CCl4 destruction rate, RRCCl4 , can be determined by the following stoichiometric reaction expression in Eq. (2). Thus, RRCCl4 is given by:
RRCCl4 ¼ RRCH4 R
M CCl4 MCH4
ð6Þ
where M denotes the molecular weight. Further, s can be expressed as
s¼
2MO2 M CH4
ð7Þ
In Eq. (2), as the R value increases the species mass fraction mCH4 and mO2 decreases, which results in a reduced reaction rate and finally the reaction will be quenched for large R values. However, the chemical kinetic aspects have not been directly considered in
Su ¼ ðaRRÞ
1=2
ð8Þ
where a denotes the thermal diffusivity. Utilizing Eq. (8) and the relationship between Su vs R, the reaction rate expression of CCl4CH4-air mixture can be obtained as a function of R and / as
ð9Þ
In Eq. (9), RRCH4 R¼0;/ represent the reaction rate of CH4 in a pure CH4-air reaction. To calculate this term, the same fast chemistry eddy breakup expression given by Eq. (9) is employed, but the role of CCl4 as both fuel and oxidizer is accounted for by using the following definition.
MC ; M CCl4 M Cl4 ; M CCl4
mfu ¼ mCH4 þ mCCl4 mOX ¼ mO2 þ mCCl4
ð10Þ
and mpr ¼ mCO2 þ mH2 O þ mHCl In the above expression, the carbon element of CCl4 is assigned in the fuel mass fraction and the total fuel mass fraction is assumed to burn just like fast chemistry reaction as CH4. A similar explanation could be made to the inclusion of Cl in CCl4. Once the CH4 reaction rate is obtained, the individual species reaction rates can be calculated by using the stoichiometric reaction expression given by Eq. (2). For example, CH4 and CCl4 reaction rate (RRCCl4 ) can be expressed as
MCH4 M CH4 þ R M CCl4 R MCH4 MCH4 þ R MCCl4
RRCH4 ¼ RRCH4 CCl4 RRCCl4 ¼ RRCH4 CCl4
ð11Þ
Further, s can be expressed as appeared in Eq. (11), that is
S¼
2M O2 þ 2RMCl2 MCH4 þ RMc
ð12Þ
M. Shin et al. / Applied Thermal Engineering 125 (2017) 20–28
In general, predictions by this model for the experimental data of the rotary kiln by Ohm [33] and dump combustor [34,35] are well agreement with the CCl4 destruction rate. In the incineration of HFCs, as mentioned earlier, the reaction rate of this mixture is believed to be determined not only by turbulence mixing but also by chemical kinetics. This is particularly true in the fuel rich and hydrogen deficient condition. However, in an incinerator operation, the excess air condition is employed in HFCs-CH4-air flames for the complete destruction of HCFs and it is possible to circumvent the effort of non-equilibrium effects in incineration process HCFs-CH4-air flames [15,16]. In Shin et al. [15], a comprehensive model and numerical calculation are made for the general and proper thermal destruction of the general waste HFCs refrigerant. In this study, HFC-23 is selected as the incineration material of HFCs material and CH4-air flame is used in an incinerator designed for CDM (Clean Development Mechanism) project. Considering the strong flame promotion effect of fluorinated hydrocarbon, however, especially in hydrogen rich and fuel lean condition, a conventional eddy breakup turbulent reaction model is employed in this excess-air incineration process. This will be briefly summarized below for the discussion of the model development of HFC incineration. The stoichiometric reaction equation of CHF3 with CH4 with auxiliary fuel is given as following;
CHF3 þ CH4 þ 2:5O2 þ 9:4N2 ! 2CO2 þ 3HF þ H2 O þ 9:4N2 þ Hfu ð13Þ where Hfu is 1074.6 kJ, the heating value of 2 mol of the fuel mixture of CHF3-CH4. In order to account for the fuel reaction rate of two species such as CHF3 and CH4, wfu in Eq. (4), the following expressions are used for the mass fractions of Eq. (4) using the reaction coefficients appearing in combustion Eq. (13).
mfu ¼ mCH4 þ mCHF3 ;
mox ¼ mO2 ;
mpr ¼ mCO2 þ mHF
ð14Þ
In Eq. (13), the component of fluorine (F) in CHF3 is assigned into the fuel part and only oxygen is considered as oxidizer fraction since HFCs shows fast chemistry behavior compared to chlorinated hydrocarbons particularly in excess air and hydrogen rich condition. Once the overall CHF3 and CH4 reaction rate is obtained, the individual CHF3 and CH4 reaction rates can be determined by the following stoichioimetric reaction expression in Eq. (13)
_ fu M CH4 =ðMCH4 þ MCHF3 Þ _ CH4 ¼ w w _ CHF3 ¼ w _ fu MCHF3 =ðM CH4 þ M CHF3 Þ w _ O2 ¼ w _ fu 2:5 MO2 =ðMCH4 þ MCHF3 Þ w _ _ wHF ¼ wfu 3 M HF =ðM CH4 þ M CHF3 Þ _ CO2 ¼ w _ fu 2 MCO2 =ðM CH4 þ M CHF3 Þ w
ð15Þ
where M denotes the molecular weight. Further, s is defined as the stoichiometric mass of oxidizer per unit mass of fuel and can be expressed as
s ¼ 2:5M O2 =ðM CH4 þ M CHF3 Þ
ð16Þ
Employing this kind of fast chemistry EBU model, Shin et al. [15] evaluate the operational data obtained from the CDM incinerator with the capacity of 70 kg/h with CH4-air flame. Numerical calculation of CHF3-CH4-air flame has been performed and evaluated successfully with the operation data of a CDM incinerator such as temperature, CHF3 destruction rate, and other species concentrations such as CO and NO at the exit of incinerator. However, not shown in this paper, the peak flame temperature occurs near exit region even 20% excess air with relatively high temperature than measurement. It is believed that the delay of the peak flame temperature toward the exit region is attributed to the improper evaluation of the fuel and oxidizer mass fraction as in Eq. (14). In summary the oxygen mass fraction is underestimated while the
23
fuel mass fraction is overestimated. Since the reaction rate in EBU model is limited by the deficient mass fraction, the underestimation of oxidizer mass fraction will cause a reduced reaction and delayed peak flame temperature with the shift of high flame temperature toward the exit. The corrected expressions of mass fraction are given in Eq. (17)
mfu ¼ mCH4 þ ðM CHF3 M CH4 Þ=M CHF3 ; mox ¼ mO2 þ ðM CHF3 MF3 Þ=MCHF3
ð17Þ
mpr ¼ mCO2 þ mHF In Shin et al. [16], a similar exercise has been made successfully for the prediction of HFC-134a thermal destruction using the fast chemistry eddy breakup model employed in Shin et al. [15]. Even if a series of research efforts of non-equilibrium turbulent modeling have been made for the halogenated hydrocarbons, in general, there is no general predictive model available to deal with halogenated hydrocarbon. Therefore, suitable modeling procedure should be continuously made specifically for each compound according to characteristics of the Damkohler number each compound in turbulent reacting flows. In this study, therefore, numerical and experimental study has been made in a tubular reactor for the development of a non-equilibrium turbulent reaction model of HFC-134a refrigerant. As mentioned in previous work [16], for the basic empiricism of incineration strategy, it is usually known that a minimum calorific value for self-sustaining flame by itself. However, in the incineration process of the halogenated compound such as HFCs, the necessity of auxiliary fuel comes from another reason in order to supply hydrogen element in order to make complete reaction by the formation of HF in Eq. (13). If we just supply only air as oxidizer without hydrogen species by the use of auxiliary fuel such as CH4, then the reaction produces fluorine molecule, F2 instead of HF, as in Eq. (13). Even if it is generally known that Freon HFC-134a can be thermally decomposed effectively only at very high temperature over 3000 °C [10,11], however, empirical findings occurring in practical calcination kiln or cokes oven have reported that if the sufficient amount of vapor exists in an high temperature environment then the refrigerant HFC-134a can be successfully decomposed as follows [16,37].
C2 H2 F4 þ 4H2 O þ 1:5O2 ! 4HF þ 3H2 O þ 2CO2
ð18Þ
The empirical work in Eq. (18) presents that for the successful conversion of C2 H2 F4 into final reaction products of HF, H2O and CO2 the supply of H and O species are required together with the necessary energy for the decomposition of C2H2F4. Therefore in this study a similar exercise has been made with the supply of O2 and water vapor in a tubular reactor to evaluate the thermal destruction of HFC-134a as a function of reactor wall temperature together with the amount of the oxidizer. For the prediction of the nonequilibrium thermal destruction of HFC-134a, a hybrid type eddy breakup model is developed by the consideration of the relative rate of the turbulent mixing and chemical reaction together. In specific, the turbulent mixing rate in Eq. (3) is obtained by eddy breakup model from Eq. (4) and the chemical reaction rate of HFC-134a is obtained from the experimental data in the tubular reactor. However, in this calculation the mass fraction of fuel, oxidizer and product are defined as following;
3 C2 H2 F4 þ O2 þ H2 O þ 3:76N2 ! 4HF þ 2CO2 þ 3:76N2 2 MC2 H2 M H2 mfu ¼ mC2 H2 F4 þ mH2 O ; M C2 H2 F4 M H2 O M F4 MO mox ¼ mO2 þ mC2 H2 F4 þ mH2 O M C2 H2 F4 MH 2 O mpr ¼ mCO2 þ mHF
ð19Þ
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M. Shin et al. / Applied Thermal Engineering 125 (2017) 20–28
The B
empirical
reaction
rate
is
assumed
as
x_ ¼ dtd ðmfu Þ ¼
ðERTa Þ
mfu mox , where mfu and mox are defined above. The preAT e exponential factor A and activation energy Ea are determined by trial and error method via numerical calculation to fit the experimental data obtained in a tubular reactor.
to satisfy the stoichiometric condition as shown above in Eq. (19). The stoichiometric amount of mass flow rates of O2 and H2O are 47.1 g/h and 17.7 g/h respectively for 100 g/h of C2H2F4. The details of experiment for the facility and measuring procedures with calibration and accuracy together with other parametric study like the change of total flow rate study are presented elsewhere [16,38].
3. Research method 3.2. Numerical model and solution procedures 3.1. Experimental facility In order to figure out the feature of the thermal decomposition of R-134a as a function of reactor wall temperature and oxygen flow rate, a series of experiments have been made using tubular reactor made of Hastelloy material as shown in Fig. 1. In order to supply the activation energy of R-134a reaction, the tubular reactor wall temperature is heated to a desired temperature. In specific, the furnace diameter and the length of reactor are 22 mm and 600 mm respectively and only 500 mm length was heated from 100 to 900 °C with 100 °C interval, while the flow rate of R-134, water vapor and N2 are fixed with 100 g/h, 17.7 g/h and 100 g/h respectively. The flow rate of oxygen and water vapor are supplied
Numerical calculation has been made by a commercial code, STAR-CCM+(Ver.9.06), which can deal with the pre-process, solver and post-process using the integrated package CAE (computer aided engineering) software in one working environment, and can solve the physical phenomenon in various fluid flow and heat transfer problems associated with chemical reaction. The same input condition was employed with the experimental condition with the flow rate of R-134a 100 g/h together with the O2 47.1 g/ h, H2O 17.7 g/h and carrier gas, N2 100 g/h via the gas mixing chamber. For completeness, the basic gas phase conservation equations for mass, momentum, energy, turbulent quantities and species concentration can be expressed as in an Eulerian framework,
Fig. 1. Detailed specification of tube reactor and boundary condition.
Table 1 General dependent variable / and diffusion coefficient C/ expression for axi-symmetric cylindrical coordinate system for the general governing Eq. (22). C2 H2 F4 þ 32 O2 þ H2 O þ 3:76N2 ! 4HF þ 2CO2 þ 3:76N2 ð22Þ. Variables
/
C/
Axial momentum
u
leff
Radial momentum
v
leff
Tangential momentum
w
leff
Kinetic energy
k
Kinetic energy dissipation rate
e
leff rk leff rs leff rh
S/
@ @x @ @x
@p 1 @ @v leff @u @x þ r @r leff r @x @x
qw @p 1 @ @v v leff @u @r þ r @r leff r @r 2leff r2 þ x @r
l
eff
r2
þ qrv þ
Gk1 qe
e j ðC 1 C k1 C 2 qeÞ
_ fu Hfu w
Specific enthalpy
h
Fuel mass fraction Oxygen mass fraction
mfu mO2
Cfu CO2
_ fu w
HF mass fraction
mHF
CHF
4
_ fu w M fu
CO2 mass fraction
mCO2
CCO2
2 Mfufu MCO2
H2O mass fraction
mH2 O
CH2 O
Mfufu MH2O
Gk1 ¼ 2leff
h @u 2 @x
þ
@ v 2 @r
h 2 i @ w2 @u @ v 2 i 2 þ vr þ r @r þ @r þ @x þ leff @w @x r
C1 = 1.44, C2 = 1.92, Cl = 0.92, rk = 0.9, re = 1.22
1 @ leff r @r
_ w
1:5 Mfufu M O2 MHF
_ w
_ w
w
2
25
M. Shin et al. / Applied Thermal Engineering 125 (2017) 20–28 !
where / represents the general dependent variable, and q and u are density and velocity vector, respectively. Further, C/ and S/ are the effective diffusion coefficients and the source term respectively for /. ! @ðq/Þ þ r ðq u /Þ ¼ r ðC/ r/Þ þ S/ @t
ð20Þ
The popular approach is the use of the two equation (k–e) model by Launder and Spalding, where a Prandtl-Kolmogorov relationship is used to correlate the turbulence viscosity, lt to the turbulent kinetic energy (k) and its dissipation rate (e) [39].
lt ¼ Cl qk2 =e
ð21Þ
As mentioned above, for the prediction of the non-equilibrium thermal destruction of HFC-134a, a hybrid type eddy breakup model is developed by the consideration of the relative rate of the turbulent mixing and chemical reaction together. To this end, empirical constants in the chemical reaction rate are determined by the trial and error method based on the experimental data in this tubular
Fig. 2. Comparison of thermal decomposition of R-134a at the reactor exit between numerical calculation and experiment as a function of wall temperature.
reactor. Further the hybrid type EBU model is employed for the prediction of the non-equilibrium turbulent reaction in this tubular reactor. The diffusion coefficient C/ and source term S/ in Eq. (20) are presented for axi-symmetric cylindrical coordinate system in Table 1 for the dependent variables such as u, v, k, e, h, mfu ; mO2 ; mHF ; mCO2 ; mH2 O respectively. The dependent variable u, v, h, mfu ; mO2 , mHF ; mCO2 , mH2 O are denoted axial and radial velocities, enthalpy, mass fractions of fuel, oxygen, hydrogen fluoride, carbon dioxide and water, respectively. In this calculation, a suitable radiation model was tested but no visible change was observed, as expected, probably due to the low emissivity by the small mean beam length caused by the small tube diameter. Fig. 2 represents a boundary condition of tube reactor for numerical calculation. 4. Results and discussion Experimental data of thermal decomposition at reactor exit as a function of temperature and oxygen concentration are presented in Tables 2 and 3, respectively. In Table 2, the fraction of the
Fig. 3. Comparison of thermal decomposition of R-134a between numerical calculations by two EBU models and experiment as a function of oxygen mole flow rate.
Table 2 Experimental data of thermal decomposition of R-134a as a function of temperature with fixed amounts of O2 and H2O. (*Carrier gas N2 100 g/h). Expt no.
Reactor wall temperature (°C)
C2H2F4 (g/h)
O2 (g/h)
H2O (g/h)
Mole ratio (C2H2F4:O2:H2O)
Decomposition percent (%) (experiment)
Decomposition percent (%) (calculation)
1 2 3 4 5 6 7 8
100 300 500 600 700 800 850 900
100 100 100 100 100 100 100 100
47.1 47.1 47.1 47.1 47.1 47.1 47.1 47.1
17.7 17.7 17.7 17.7 17.7 17.7 17.7 17.7
1:1.5:1 1:1.5:1 1:1.5:1 1:1.5:1 1:1.5:1 1:1.5:1 1:1.5:1 1:1.5:1
0 0 0 1.1 17.7 79.6 100 100
0 0 0 5.2 20.7 80.4 100 100
Table 3 Experimental and numerical data of the thermal decomposition of R-134a as a function of oxygen flow using two EBU models (*Carrier gas N2 100 g/h and 900 °C wall temperature and the stoichiometric amount of H2O.) Expt no.
Reactor wall temperature (°C)
C2H2F4 (g/h)
O2 (g/h)
H2O (g/h)
Mole ratio (C2H2F4:O2:H2O)
Experimental decomposition percent (%)
Numerical decomposition percent (%) by fast chemistry EBU model
Numerical Decomposition percent (%) by hybrid model
8 9 10 11 12 13 14
900 900 900 900 900 900 900
100 100 100 100 100 100 100
47.1 0 9.4 15.7 21.9 28.3 37.7
17.7 17.7 17.7 17.7 17.7 17.7 17.7
1:1.5:1 1:0.0:1 1:0.3:1 1:0.5:1 1:0.7:1 1:0.9:1 1:1.2:1
100 25 50.8 – – 94 100
100 31.8 53.4 67.4 77.6 86.8 94.6
100 22.7 57.1 75.7 86.2 94.7 100
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HFC-134a thermal decomposition is presented with the wall temperature increase from 100 °C to 900 °C for the fixed value of the stoichiometric amount of oxygen and water vapor. In Table 3, the thermal decomposition data are presented as a function of oxygen concentration with the fixed value of wall temperature and 900 °C and stoichiometric amount of H2O concentration. As shown in Table 2, no thermal decomposition of HFC 134a occurs up to the wall temperature of 500 °C and only very small amount of decomposition, that is, 1.1% occurs at temperature 600 °C even if there exist stoichiometric amount of oxygen and water vapor. But the fraction of thermal decomposition increases 17.7, 79.6 and 100% as wall temperature increases 700, 800, and 850 °C, respectively. This data suggests that at least more than 600 °C is necessary in order to react HFC-134a with O2 and water vapor. Using this
decomposition data at the reactor exit, empirical constants of pre-exponential factor (A), empirical constant B and activation Ea
energy (Ea) in the chemical reaction rate, that is, k ¼ AT B eðRT Þ was determined by trial and error method in order to fit the destruction fraction of HFC-134a in Table 2 using the hybrid type EBU model for the development of the non-equilibrium turbulent reaction model of HFC-134a incineration. The value of A, b and Ea obtained are A = 1.0, B = 0.38, Ea = 8.8 107 J/kmol, respectively. The calculated thermal decomposition data are well compared with experimental data and presented in Fig. 2 and Table 2, respectively. Further, using the hybrid EBU model with empirical constants determined above, the thermal destruction data as in Table 3 are predicted as a function of oxygen concentration with fixed wall
(a) Velocity vector( inlet velocity,0.126 m/s)
(b) Temperature distribution
(c) Mass fraction of HFC-134a Fig. 4. Calculation results for the wall temperature of 900 °C and stoichiometric condition of oxygen and water vapor by conventional EBU model for the Exp. No. 8 in Table 2 or 3.
(a) Velocity vector( inlet velocity,0.126 m/s)
(b) Temperature distribution
(c) Mass fraction of HFC-134a Fig. 5. Calculation results for the wall temperature of 900 °C and stoichiometric condition of oxygen and water vapor by the non-equilibrium EBU model for the Exp. No. 8 in Table 2 or 3.
M. Shin et al. / Applied Thermal Engineering 125 (2017) 20–28
27
Fig. 6. Comparison of numerical calculations by two EBU models along the axial distance of tube reactor.
temperature 900 °C. In this calculation, the fast chemistry EBU model used in Shin et al. [16] is also employed for the comparison purpose. Comparison between calculations by two EBU models and experiment has been made and presented at Table 3 and in Fig. 3 for the thermal decomposition of R-134a as a function of O2 concentration. In general, the prediction by hybrid-type EBU model agrees better than the conventional fast chemistry model without any non-equilibrium incorporation. Since the comparison has been made only for the concentration of HFC-134a at the exit, the calculated profiles are presented in detail in order to see the difference of two models. In Figs. 4 and 5, the detailed profiles of HFC-134a concentration, temperature and velocity field are presented respectively. Even if the overall thermal decomposition fraction by two models look similar at the exit, the profiles of temperature and HFC-134a concentration differ significantly. As shown in Fig. 5, the reaction of HFC-134a occurs very slowly at low temperature entrance region due the non-equilibrium effect. Further, detailed profile comparisons of temperature and HFC-134a have been made along the reactor centerline axis using two models as shown in Fig. 6. This clearly shows the non-equilibrium effect of turbulence reaction model plays a significant role especially in the significant region of the non-equilibrium chemistry effect especially in low flame temperature region.
5. Conclusion Considering the global warming potential of HFC-134a (C2H2F4) with the substantial generation of this refrigerant as waste material in various industrial sectors, the development of proper thermal destruction method of HFC-134a is of great practical significance especially in the generation of fluorine material like HF and CaF2. Experiment and numerical calculation have made for a tubular type in order to figure out the important feature of the thermal decomposition of HFC-134a as a function of temperature and reactants such as oxygen and water vapor. Detailed review and discussions are made for the modeling and nonequilibrium effect in the incineration process of various halogenated hydrocarbons using a couple of modified versions of eddy breakup (EBU) model developed from the original EBU model. Employing this non-equilibrium hybrid EBU model, the calculated overall destruction rate of HFC-134a was in good agreement with the experimental data obtained for the same reactor. But the detailed HFC-134a concentration profiles between the fast chemistry and non-equilibrium EBU models show a significant difference even if the calculated destruction rates of HFC-134a appear similar at the exit. In near future, numerical calculation will be
made for the practical incinerator of HFC-134a in a practical incinerator using the model developed in this study. Acknowledgement This project is supported by the ‘‘R&D Center for reduction of Non-CO2 Greenhouse gases (2013001690006)” funded by Korea Ministry of Environment (MOE) as ‘‘Global Top Environment R&D Program”. References [1] China Fluoro Technology Co., CDM Monitoring Report, UNFCCC (United Nations Framework Convention on Climate Change), CDM Reference Number 1194, 2007. [2] D.S. Jang, M.S. Shin, Y.G. Lee, Y.J. Kim, On a thermal destruction of HFC-134a and HFC-23 by CH4 and water electrolysis gas in a CDM incinerator, in: The 2nd 3R International Scientific Conference on Material Cycles and Waste Management, 2015, pp. 205–209. [3] US Environmental Protection Agency, Complying with the Section 608 Refrigerant Recycling Rule, 2011.. [4] A. Jones, H. Bliss, C. Walker, Rates of hydrogen chloride oxidation, AIChE J. 12 (2) (1996) 260–265. [5] J.J. Santoleri, Chlorinated hydrocarbon waste disposal and recovery systems, Chem. Eng. Prog. 69 (1) (1973) 66–74. [6] J.M. Robinson, Experimental Shooting Limits for Methane-Chlorinated Hydrocarbon Flames M.S. Thesis, Massachusetts of Technology, Department of Chemical Engineering, 1981. [7] F.H. Garner, R. Long, A.J. Graham, A. Badakshan, Effect of certain halogenated methanes on premixed and diffusion flames, Symp. Int. Combust. 6 (1) (1956) 802–806. [8] H.A. Valerias, Burning Velocities and Rates of Methane Chlorinated Hydrocarbon Flames, Master Thesis, Chemical Engineering Dept, Massachusetts Institute of Technology, 1982. [9] W.A. Rosser, H. Wise, J. Miller, Mechanism of combustion inhibition by compounds containing halogens, Symp. Int. Combust. 7 (1) (1959) 175–182. [10] C. Womeldorf, W. Grosshandler, M. Mitchell, Selection of a simulant of CF3Br for use in engine nacelle certification tests, Halon Opt. Tech. Work. Conf. 5 (1995) 197–210. [11] V.R. Katta, W. Roquemore, Effect of nitromethane on soot formation in heptane jet diffusion flame, in: Joint Meeting – US Sections of the Combustion Insti, vol. 4, 2005. [12] G.W. Gmurczyk, W.L. Grosshandler, D.L. Lowe, Suppression effectiveness of extinguishing agents under highly dynamic conditions, Proc. Int. Symp. Fire Saf. Sci. 4 (1995) 925–936. [13] V.V. Azatyan, Y.N. Shebeko, A.Y. Shebeko, Effect of self-inhibition at a flame propagation in rich gaseous mixtures of combustible-air-diluent, J. Loss Prevent. Process Ind. 20 (4–6) (2007) 486–493. [14] V.V. Azatyan, I.A. Bolodyan, V.Y. Navtsenya, The dependence of the rules governing gas-phase combustion on the competition between chain propagation and chain termination reactions, Russ. J. Phys. Chem. A 84 (3) (2010) 383–390. [15] M.S. Shin, D.S. Jang, J. Ha, Experimental and numerical calculation of waste R134a thermal decomposition for energy and fluorine material recovery, J. Mater. Cycl. Waste Manage. 18 (3) (2015) 399–406. [16] L. Zhang, S. Li, Theoretical study of the kinetics for the hydrogen abstraction of 1,1,1,2-tetrafluoroethane (HFC134a) by hydroxyl radical, J. Mol. Struct.: THERMOCHEM 869 (2008) 6–10.
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[17] V.A. Cundy et al., Rotary kiln incineration I. An in-depth study – liquid injection, J. Air Pollut. Control Assoc. 39 (1) (1989) 63–75. [18] J. Lee, J. Shim, Treatment of Waste Refrigerants by Calcimation Reactor, Korea Patent No. 10-2013-0002125, 2013. [19] W.D. Clark, M.P. Heap, W. Richter, W.R. Seeker, The prediction of liquid injection hazardous waste incineration, in: ASME/AIChE National Heat Transfer Conference 84-HT-13, 1984. [20] Y. Yang, J.A. Marc, M.A. Reuter, Modeling of the fuel stream and combustion in a rotary-kiln hazardous waste incinerator, in: Third Int’l Conference on CFD in the Minerals and Process Industries, 2003, pp. 10–12. [21] R.W. Bilger, Turbulent jet diffusion flames, Energy Combust. Sci. 1 (1976) 87–109. [22] D.B. Spalding, Mixing and chemical reaction in steady confined turbulent flames, Symp. Int. Combust. 13 (1) (1971) 649–657. [23] B.F. Magnussen, H. Hjertager, On mathematical modeling of turbulent combustion with special emphasis on soot formation and combustion, Symp. Int. Combust. 16 (1) (1976) 719–729. [24] R.W. Bilger, Perturbation analysis of turbulent non-premixed combustion, Combust. Sci. Technol. 22 (5–6) (1980) 251–261. [25] N. Peters, F.A. Williams, Lift-off characteristics of turbulent jet diffusion flames, AIAA J. 21 (3) (1983) 423–429. [26] A. Chakravarty, F.C. Lockwood, O. Sinicropi, The prediction of burner stability limits, Combust. Sci. Technol. 42 (1–2) (1984) 67–86. [27] D.S. Jang, S. Acharya, Calculation of non-equilibrium effects in turbulent nonpremixed combustion, in: ASME/AIChE National Heat Transfer Conference 87-HT-52, 1987. [28] D.S. Jang, S. Acharya, D. West, L. Herbert, A moment closure method for modeling multistep chlorination reactions, in: AIChE Spring Meeting AIChE 91/ 33a, 1991.
[29] D.S. Jang, S. Acharya, Two dimensional modeling of waste combustion in a kiln, in: AIChE Spring Meeting, 1988. [30] F.C. Lockwood, A.P. Salooja, S.A. Syed, A prediction method for coal-fired furnace, Combust. Flame 38 (1980) 1–15. [31] M.S. Shin, Y.J. Kim, D.S. Jang, Y.G. Lee, K.W. Lee, Numerical study on the cofiring flame characteristics of CH4 and refrigerant R-100(CCl4), J. Kor. Soc. Waste Manage. 31 (8) (2014) 868–869. [32] D.S. Jang, S. Acharya, A numerical simulation of CCl4 destruction in a twodimensional kiln, Energy R&D J. Kor. Res. Inst. Energy 12 (3) (1990) 94–111. [33] T.I. Ohm, D.S. Jang, J.O. Chae, A study for the advanced design of rotary kiln incinerator III: 3-dimensional CCl4/CH4 gas-phase turbulent reaction model, Kor. Soc. Energy Eng. 2(1) (1993) 54–67. [34] Y.N. Chun, K.J. Lee, H.O. Song, A numerical simulation of hazardous waste destruction in a three-dimensional dump incinerator, Kor. J. Chem. Eng. 19 (1) (2002) 20–27. [35] Y.N. Chun, O.J. Jung, S.W. Kim, H.O. Song, Numerical and experimental studies of CCl4 destruction in a dump incinerator, Environ. Technol. 24 (2) (2003) 131–142. [36] I. Glassman, Combustion, Academic Press, New York, 1977. [37] J. Lee, J. Shim, Treatment of Waste Refrigerants by Calcination Reactor, Korea Patent No. 10-2013-0002125, 2013. [38] J.S. Ha, 2014 Annual Report for the Development of Waste Refrigerant Separation/Refinement, Reclamation and Thermal Treatment Presented for the R&D Center for the Reduction of Non-CO, 2 Green House Gases, Green Chemistry & Engineering, Korea Research Institute of Chemical Technology, 2015. [39] B.E. Launder, D.B. Spalding, Mathematical Models of Turbulence, Academic Press, London/New York, 1972.
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Applied Thermal Engineering journal homepage: www.elsevier.com/locate/apthermeng
Research Paper
Development of a non-equilibrium turbulent reaction model of HFC134a in a tubular reactor Misoo Shin a, Dongsoon Jang a,⇑, Jongwook Ha b, S. Acharya c a
Department of Environmental Engineering, Chungnam National University, 99 Daehak-ro, Yuseong-gu, Daejeon 34134, South Korea Korea Research Institute of Chemical Technology, 141 Gajeong-ro, Yusunggu, Daejeon 305-600, South Korea c Department of Mechanical Engineering, Louisiana State University, Baton Rouge, LA 70803, USA b
h i g h l i g h t s A hybrid-type EBU model is developed for HFC-134a incineration. Experiment and numerical prediction data are well compared each other. The hybrid EBU model shows big difference with the conventional model.
a r t i c l e
i n f o
Article history: Received 18 October 2016 Revised 23 May 2017 Accepted 20 June 2017 Available online 22 June 2017 Keywords: Halogen compound incineration HFC-134a Non-equilibrium turbulent reaction model Hybrid-type eddy breakup model
a b s t r a c t Detailed review and discussions are made for the modeling and non-equilibrium effect in the incineration process of various halogenated hydrocarbons using a couple of modified versions of eddy breakup (EBU) model developed from the original EBU model by Magnussen and Hjertager. Further experimental and numerical studies are made for the development of a hybrid-type EBU model in order to evaluate the non-equilibrium effect of HFC-134a reaction in a tubular reactor. For the numerical calculation using the hybrid-type EBU model, especially, the relative rate of the turbulent mixing and chemical reaction should be incorporated in a harmonic mean expression. Thus the overall reaction rate expression of HFC-134a reaction is determined by the fitting the empirical constants with experimental data obtained in the tubular reactor via trial and error method. Thereby, the empirical constants in the equation of the E _ ¼ kC C2 H2 F4 C H2 O C O2 ¼ AT B eRTa C C2 H2 F4 C H2 O C O2 was determined, Arrehenius type reaction rate, that is, x where A = 1.0, B = 0.38, Ea = 8.8 107 J/kmol, respectively. Employing this non-equilibrium model, the calculated overall destruction rate of HFC-134a was in good agreement with the experimental data obtained for the same reactor. Further, the comparison of HFC-134a concentration profiles between the fast chemistry and non-equilibrium EBU models show a significant difference even if the calculated destruction rates of HFC-134a appear similar at the exit. In near future, numerical calculation will be made for the practical incinerator of HFC-134a in a practical incinerator using the model developed in this study. Ó 2017 Elsevier Ltd. All rights reserved.
1. Introduction These days, reclaim and recovery services of used refrigerants such as HFCs are one of major concern in order to minimize the environmental impact of waste refrigerants especially in the aspect of the release of the material of high global warming potential. However, a significant part of recovered refrigerants have to be destroyed inevitably in a practical incinerator in an environmentally friendly manner when their compositions or purity are inferior than specified condition for reclamation [1]. As a matter ⇑ Corresponding author. E-mail address: [email protected] (D. Jang). http://dx.doi.org/10.1016/j.applthermaleng.2017.06.105 1359-4311/Ó 2017 Elsevier Ltd. All rights reserved.
of fact, the incineration has been known over several decades as a preferred method for hazardous waste disposal compared to other conventional methods because of its potential for immediate detoxification and the potential for energy recovery. A proper incineration method for the waste HFCs (hydroflurocarbons) has become a viable method not only for the reduction of the release of the greenhouse gases but also for the recovery of the valuable material such as CaF2 (calcium fluoride) [2]. However, it is generally known the practical incineration process of most refrigerants containing halogen species such as F and Cl could not be properly modeled by typical fast chemistry turbulent reaction model due to not only its flame inhibition effect but also the low heating value of halogen containing species. Thus a proper non-equilibrium
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turbulent reaction model is quite required for a practical incinerator to develop a reliable predictive model in order to obtain proper design and operational conditions for the practical HFCs incineration in turbulent reacting flows. To this end, in this study detailed review and discussions are made first for the modeling and nonequilibrium effect in the incineration process of various halogenated hydrocarbons. And a numerical and experimental study has been made in a tubular reactor for the development of proper non-equilibrium turbulent reaction model of HFC-134a.
2. Review of halogenated compound incineration and numerical modeling Research on the combustion of various halogenated hydrocarbons has been done mainly in incineration area of hazardous waste, for chlorinated hydrocarbons, generated from chemical industries since 1970. Since incineration is known as a preferred tool for hazardous waste disposal compared to other conventional methods, this technology is also certified by United Nations Framework Convention on Climate Change (UNFCC) for the thermal destruction of halogenated hydrocarbons. However, as well known, only general guideline is given to accomplish 99.99% thermal destruction efficiency (DRE), based on the empiricism, suggesting that it requires flame temperature higher than 1200 °C and the residence time more than 2 s [3]. Since halogenated hydrocarbons not only serve as fire extinguishing agents, but also some of those cannot produce self-sustaining flame. One of the typical ways for proper incineration, therefore, is to co-burning of this compound with auxiliary fuels of containing many hydrogen atoms such as CH4 and H2. Halogenated hydrocarbons with the aid of this kind of fuel can be effectively decomposed by the reaction with hydrogen and oxygen; hydrogen combines with halogen to form HCl or HF. In this process, it is noted that the halogen atom originated from halogenated hydrocarbon plays not only the role of oxidizer but also inhibition action through the combination of the hydrogen containing in auxiliary fuel. In this process, the conventional strategy of introducing excess air to ensure complete combustion of the waste stream in practical incinerators is not quite desirable to the incineration of CHCs (Chlorinated Hydrocarbons) because of the following Deacon reaction [4].
1 2HCl þ O2 ! Cl2 þ H2 O 2
ð1Þ
Excess oxygen can lead to pure chlorine gas, Cl2, which is more toxic and difficult to scrub out of stack gases than HCl [5]. This reaction together with the inhibition effect makes incineration of CHCs difficult to accomplish the goal of DRE 99.99%. Similar argument can be applied to the HFCs refrigerants containing fluorine. In general, the fundamental studies have shown that the incineration process of halogenated hydrocarbons is usually inhibited by the reaction of halogen molecule (X2) with hydrogen H especially in preheating region since the activation energy of inhibition reaction is much lower than that of the chain branching reaction [6]. Garner et al. [7] and Valeiras [8] showed that the order of effectiveness of in inhibition for chloromethane by decreasing flame velocity is CCl4, CHCl3, CH2Cl2 and CH3Cl with the increase of the chlorine molecule. The flame velocities of the mixture are, of course, lower than methane and decrease with the increase of the molar fraction of chlorine compound to the auxiliary fuel (R), for example R = mole of CHCs/mole of CH4 as in Eq. (2).
CH4 þ RCCl4 þ 2O2 þ 7:25N2 ! ð1 þ RÞCO2 þ 2ð1 RÞH2 O þ 4RHCl þ 7:25N2
ð2Þ
Further as the chlorine content increases by R or by the change compound, the flammability range decreases and the maximum flame velocity shifts from fuel rich to fuel lean condition [9]. The net effect of the halogen inhibition on flame is to increase the length of the preheat zone over that of the uninhibited flame but the higher temperature in the reaction zone increases exponentially the chain branching reaction rate enough to overcome the rates of the inhibition reaction [7]. Further, iodine and bromine compounds are found to be more strong inhibitors than chlorine and fluorine compounds [6]. Even if HFCs have been developed as essential substitute for the ozone depleting refrigerants or sprays such as CFCs and HCFCs containing chlorine compounds, some of the alternative HFCs show unexpectedly high flammability. Therefore, in order to ensure fire safety application of these substances, there has been a lot of fundamental combustion study about fluorinated hydrocarbons [8–14]. A detailed reviews about this have been made elsewhere [15]. A series of detailed chemistry data of HFC-134a appear only in the research area of degradation pathways into trifluoroacetic acid (TFA) in the atmosphere instead of combustion field like waste incineration. In detail, the OH radical is expected to be the most important oxidation of HFCs degradation in daytime troposphere. Extensive experimental studies have been taken on the kinetics of the hydrogen abstraction of HFC134a by hydroxyl radical since this reaction plays an important role in the gas-phase reaction of HFC-134a in the atmosphere [16]. The rate constants determined are presented but unfortunately the temperature ranges are below than 500 K, which is far below the flame temperature. Along with the fundamental combustion study, even though a lot of research efforts [17–19] have been made about thermal destruction of halogenated compound both in large scale incinerators and also model developments have been made to predict the performance of incineration processes, these studies do not provide a predictive tool for resolution of the detailed distributions of the turbulent flow field. Even though Yang et al. [20] made a detailed numerical calculation for a practical rotary kiln incinerator, they did not include any halogenated compounds as principal organic hazardous constituents. In general, there are two fundamental steps for the process of turbulent combustion; the first step is turbulent mixing between fuel and air, and the next chemical reaction by the molecular collision between fuel and oxidizer. In order to make complete combustion successfully, therefore, it is necessary to mix efficiently fuel with air by eddy breakup action and then fuel molecule will react with oxygen molecule for chemical reaction. The overall reaction rate occurring in a series of competitive process can be typically expressed empirically in a harmonic form such as
Ov erall reaction rate
1 1 turbulent mixing rate
1 þ chemistry
ð3Þ rate
The basic problem of the gaseous turbulent combustion lies in the proper modeling of the well-known turbulent correlation term, which arises in the non-linear chemical reaction source term like Reynolds stress in momentum equation. Most turbulent combustion models, however, detour the difficulty by ignoring the chemical kinetic aspects of the problem as shown in the harmonic mean expression of overall reaction rate in Eq. (3). In detail, the chemistry rate is fast and thereby very large compared to turbulent mixing rate. Therefore the term can be neglected. Using the assumption of fast chemistry, the turbulent mixing will be governed only by the limit of the mixing rate in Eq. (3). Although mixing limited turbulence only models have been shown to be successful tools in predicting the overall profiles of main species, serious drawback exist, since the conventional fast chemistry turbulent reaction models such as the conserved scalar method of Bil-
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ger [21], the eddy breakup models of Spalding [22] and Magnussen and Hjertager [23] do not describe the non-equilibrium characteristics of halogen compounds. Since the effect of inhibitive chemistry reaction is as important as that of turbulent mixing an appropriate turbulent combustion model had to be used, where chemical kinetics as well as turbulence mixing plays an important role. In the past, considerable progress has been made in the modeling of chemical kinetic processes in turbulent flows for various areas such as flame extinction in strong shear flows, NOx generation in pollutant formation and chlorination process in chemical reactor [24–28]. However, general non-equilibrium turbulent reaction model is not yet available especially for the halogenated incineration processes. For the numerical modeling efforts of CHCs, CCl4 destruction has been done mainly since CCl4 was one of the major hazardous waste to be incinerated. Therefore a number of investigators have focused on the modeling of the turbulent reaction in a premixed CCl4-CH4-air flame. As emphasized above, the reaction rate of this mixture is believed to be determined not only by turbulence mixing but also by chemical kinetics. This is particularly, true since CCl4 is incapable of a self-sustaining flame due to inhibition feature as well as the low enthalpy of combustion. In order to resolve the non-equilibrium effects in CCl4-CH4-air flames, detailed chemical kinetic data is required. And a number of research efforts have been directed on this end [17], but there was no wellestablished chemical kinetic data available. Thus an empirical modeling approach was suggested in the CCl4 incineration process. The first reduced reaction model for CCl4 destruction [29] is presented briefly in the following for the discussion of the modified version of EBU model. The reaction rate of CH4 (RRCH4 ) in a CCl4CH4-air mixture is given by the phenomenological eddy breakup model proposed by Magnussen and Hjertager [23], i.e.
this turbulent reaction model. Shin et al. [31] employs a similar eddy breakup model for the reaction of the mixture of CH4-CCl4. But they consider the mixture of CH4-CH3Cl behaves together as fast chemistry reaction, while Jang and Acharya [29] assume that only CH4 reaction is governed by fast chemistry. The reason why Shin et al. employ the fast chemistry model for the total mixture of CH4-CCl4 was in that the small amount of CCl4 would not play ant significant inhibition action most over strong turbulent reaction region. This model is reasonable only for the case of the small amount of CCl4 presence. Further, Jang and Acharya [32] propose a non-equilibrium turbulent reaction model using burning velocity data given by Valeiras [8] for CH4-CCl4-air mixture e in 2-D hypothetical kiln. That is, a modified fast chemistry turbulent reaction model was developed to describe the flame inhibition due to the presence of CCl4, considering the corresponding burning velocity data of these mixtures. This modified EBU model was employed with reasonably good prediction for numerical calculation of practical incinerators for a rotary kiln and dump combustor [33–35]. The experimental burning velocity data of premixed CCl4-CH4-air mixture as a function of R and equivalence ratio in a laboratory Bunsen flame study [6] is utilized to obtain a more realistic reaction rate expression. In the aspect to review the modified eddy breakup model for non-equilibrium turbulent combustion of halogenated hydrocarbons, let us briefly summarize this model. Based on the thermal theory of flame propagation [36], the reaction rate (RR) can be related to the flame burning velocity (Su ) by
pr m m Am fu =k; q A0 OX ð=kÞ; qA0 RRCH4 ¼ minimum of q s ð1 þ sÞ k
2 Su a R; / RRCH4 R¼0;/ RRCH4CCl4 at R;/ ¼ 2 Su a R ¼ 0; /
ð4Þ where s is the stoichiometric oxidizer mass per unit mass of fuel and A0 and A are empirical constants given by Lockwood et al. [30]. The basic idea of this model is based on premise that only the reaction of CH4 in a CCl4-CH4-air mixture is governed by the fast chemistry concept and the destruction of CCl4 can be described by the stoichiometric reaction Eq. (2) according to the destruction of CH4. It is generally expected that the CH4 reaction rate given by Eq. (4) will be smaller compared to that in a pure CH4-airmixture, (i.e. R = 0.0) because of the decreased mass fractions due the presence of CCl4 and HCl compound. In order to account for the CH4 reaction rate in CCl4-CH4-air mixture, the following expressions arc used for the mass fractions appearing in Eq. (4)
mfu ¼ mCH4 ; mOX ¼ mO2
and mpr ¼ mCO2 þ mH2 O þ mHCl
ð5Þ
Once the CH4 reaction rate is obtained, the CCl4 destruction rate, RRCCl4 , can be determined by the following stoichiometric reaction expression in Eq. (2). Thus, RRCCl4 is given by:
RRCCl4 ¼ RRCH4 R
M CCl4 MCH4
ð6Þ
where M denotes the molecular weight. Further, s can be expressed as
s¼
2MO2 M CH4
ð7Þ
In Eq. (2), as the R value increases the species mass fraction mCH4 and mO2 decreases, which results in a reduced reaction rate and finally the reaction will be quenched for large R values. However, the chemical kinetic aspects have not been directly considered in
Su ¼ ðaRRÞ
1=2
ð8Þ
where a denotes the thermal diffusivity. Utilizing Eq. (8) and the relationship between Su vs R, the reaction rate expression of CCl4CH4-air mixture can be obtained as a function of R and / as
ð9Þ
In Eq. (9), RRCH4 R¼0;/ represent the reaction rate of CH4 in a pure CH4-air reaction. To calculate this term, the same fast chemistry eddy breakup expression given by Eq. (9) is employed, but the role of CCl4 as both fuel and oxidizer is accounted for by using the following definition.
MC ; M CCl4 M Cl4 ; M CCl4
mfu ¼ mCH4 þ mCCl4 mOX ¼ mO2 þ mCCl4
ð10Þ
and mpr ¼ mCO2 þ mH2 O þ mHCl In the above expression, the carbon element of CCl4 is assigned in the fuel mass fraction and the total fuel mass fraction is assumed to burn just like fast chemistry reaction as CH4. A similar explanation could be made to the inclusion of Cl in CCl4. Once the CH4 reaction rate is obtained, the individual species reaction rates can be calculated by using the stoichiometric reaction expression given by Eq. (2). For example, CH4 and CCl4 reaction rate (RRCCl4 ) can be expressed as
MCH4 M CH4 þ R M CCl4 R MCH4 MCH4 þ R MCCl4
RRCH4 ¼ RRCH4 CCl4 RRCCl4 ¼ RRCH4 CCl4
ð11Þ
Further, s can be expressed as appeared in Eq. (11), that is
S¼
2M O2 þ 2RMCl2 MCH4 þ RMc
ð12Þ
M. Shin et al. / Applied Thermal Engineering 125 (2017) 20–28
In general, predictions by this model for the experimental data of the rotary kiln by Ohm [33] and dump combustor [34,35] are well agreement with the CCl4 destruction rate. In the incineration of HFCs, as mentioned earlier, the reaction rate of this mixture is believed to be determined not only by turbulence mixing but also by chemical kinetics. This is particularly true in the fuel rich and hydrogen deficient condition. However, in an incinerator operation, the excess air condition is employed in HFCs-CH4-air flames for the complete destruction of HCFs and it is possible to circumvent the effort of non-equilibrium effects in incineration process HCFs-CH4-air flames [15,16]. In Shin et al. [15], a comprehensive model and numerical calculation are made for the general and proper thermal destruction of the general waste HFCs refrigerant. In this study, HFC-23 is selected as the incineration material of HFCs material and CH4-air flame is used in an incinerator designed for CDM (Clean Development Mechanism) project. Considering the strong flame promotion effect of fluorinated hydrocarbon, however, especially in hydrogen rich and fuel lean condition, a conventional eddy breakup turbulent reaction model is employed in this excess-air incineration process. This will be briefly summarized below for the discussion of the model development of HFC incineration. The stoichiometric reaction equation of CHF3 with CH4 with auxiliary fuel is given as following;
CHF3 þ CH4 þ 2:5O2 þ 9:4N2 ! 2CO2 þ 3HF þ H2 O þ 9:4N2 þ Hfu ð13Þ where Hfu is 1074.6 kJ, the heating value of 2 mol of the fuel mixture of CHF3-CH4. In order to account for the fuel reaction rate of two species such as CHF3 and CH4, wfu in Eq. (4), the following expressions are used for the mass fractions of Eq. (4) using the reaction coefficients appearing in combustion Eq. (13).
mfu ¼ mCH4 þ mCHF3 ;
mox ¼ mO2 ;
mpr ¼ mCO2 þ mHF
ð14Þ
In Eq. (13), the component of fluorine (F) in CHF3 is assigned into the fuel part and only oxygen is considered as oxidizer fraction since HFCs shows fast chemistry behavior compared to chlorinated hydrocarbons particularly in excess air and hydrogen rich condition. Once the overall CHF3 and CH4 reaction rate is obtained, the individual CHF3 and CH4 reaction rates can be determined by the following stoichioimetric reaction expression in Eq. (13)
_ fu M CH4 =ðMCH4 þ MCHF3 Þ _ CH4 ¼ w w _ CHF3 ¼ w _ fu MCHF3 =ðM CH4 þ M CHF3 Þ w _ O2 ¼ w _ fu 2:5 MO2 =ðMCH4 þ MCHF3 Þ w _ _ wHF ¼ wfu 3 M HF =ðM CH4 þ M CHF3 Þ _ CO2 ¼ w _ fu 2 MCO2 =ðM CH4 þ M CHF3 Þ w
ð15Þ
where M denotes the molecular weight. Further, s is defined as the stoichiometric mass of oxidizer per unit mass of fuel and can be expressed as
s ¼ 2:5M O2 =ðM CH4 þ M CHF3 Þ
ð16Þ
Employing this kind of fast chemistry EBU model, Shin et al. [15] evaluate the operational data obtained from the CDM incinerator with the capacity of 70 kg/h with CH4-air flame. Numerical calculation of CHF3-CH4-air flame has been performed and evaluated successfully with the operation data of a CDM incinerator such as temperature, CHF3 destruction rate, and other species concentrations such as CO and NO at the exit of incinerator. However, not shown in this paper, the peak flame temperature occurs near exit region even 20% excess air with relatively high temperature than measurement. It is believed that the delay of the peak flame temperature toward the exit region is attributed to the improper evaluation of the fuel and oxidizer mass fraction as in Eq. (14). In summary the oxygen mass fraction is underestimated while the
23
fuel mass fraction is overestimated. Since the reaction rate in EBU model is limited by the deficient mass fraction, the underestimation of oxidizer mass fraction will cause a reduced reaction and delayed peak flame temperature with the shift of high flame temperature toward the exit. The corrected expressions of mass fraction are given in Eq. (17)
mfu ¼ mCH4 þ ðM CHF3 M CH4 Þ=M CHF3 ; mox ¼ mO2 þ ðM CHF3 MF3 Þ=MCHF3
ð17Þ
mpr ¼ mCO2 þ mHF In Shin et al. [16], a similar exercise has been made successfully for the prediction of HFC-134a thermal destruction using the fast chemistry eddy breakup model employed in Shin et al. [15]. Even if a series of research efforts of non-equilibrium turbulent modeling have been made for the halogenated hydrocarbons, in general, there is no general predictive model available to deal with halogenated hydrocarbon. Therefore, suitable modeling procedure should be continuously made specifically for each compound according to characteristics of the Damkohler number each compound in turbulent reacting flows. In this study, therefore, numerical and experimental study has been made in a tubular reactor for the development of a non-equilibrium turbulent reaction model of HFC-134a refrigerant. As mentioned in previous work [16], for the basic empiricism of incineration strategy, it is usually known that a minimum calorific value for self-sustaining flame by itself. However, in the incineration process of the halogenated compound such as HFCs, the necessity of auxiliary fuel comes from another reason in order to supply hydrogen element in order to make complete reaction by the formation of HF in Eq. (13). If we just supply only air as oxidizer without hydrogen species by the use of auxiliary fuel such as CH4, then the reaction produces fluorine molecule, F2 instead of HF, as in Eq. (13). Even if it is generally known that Freon HFC-134a can be thermally decomposed effectively only at very high temperature over 3000 °C [10,11], however, empirical findings occurring in practical calcination kiln or cokes oven have reported that if the sufficient amount of vapor exists in an high temperature environment then the refrigerant HFC-134a can be successfully decomposed as follows [16,37].
C2 H2 F4 þ 4H2 O þ 1:5O2 ! 4HF þ 3H2 O þ 2CO2
ð18Þ
The empirical work in Eq. (18) presents that for the successful conversion of C2 H2 F4 into final reaction products of HF, H2O and CO2 the supply of H and O species are required together with the necessary energy for the decomposition of C2H2F4. Therefore in this study a similar exercise has been made with the supply of O2 and water vapor in a tubular reactor to evaluate the thermal destruction of HFC-134a as a function of reactor wall temperature together with the amount of the oxidizer. For the prediction of the nonequilibrium thermal destruction of HFC-134a, a hybrid type eddy breakup model is developed by the consideration of the relative rate of the turbulent mixing and chemical reaction together. In specific, the turbulent mixing rate in Eq. (3) is obtained by eddy breakup model from Eq. (4) and the chemical reaction rate of HFC-134a is obtained from the experimental data in the tubular reactor. However, in this calculation the mass fraction of fuel, oxidizer and product are defined as following;
3 C2 H2 F4 þ O2 þ H2 O þ 3:76N2 ! 4HF þ 2CO2 þ 3:76N2 2 MC2 H2 M H2 mfu ¼ mC2 H2 F4 þ mH2 O ; M C2 H2 F4 M H2 O M F4 MO mox ¼ mO2 þ mC2 H2 F4 þ mH2 O M C2 H2 F4 MH 2 O mpr ¼ mCO2 þ mHF
ð19Þ
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M. Shin et al. / Applied Thermal Engineering 125 (2017) 20–28
The B
empirical
reaction
rate
is
assumed
as
x_ ¼ dtd ðmfu Þ ¼
ðERTa Þ
mfu mox , where mfu and mox are defined above. The preAT e exponential factor A and activation energy Ea are determined by trial and error method via numerical calculation to fit the experimental data obtained in a tubular reactor.
to satisfy the stoichiometric condition as shown above in Eq. (19). The stoichiometric amount of mass flow rates of O2 and H2O are 47.1 g/h and 17.7 g/h respectively for 100 g/h of C2H2F4. The details of experiment for the facility and measuring procedures with calibration and accuracy together with other parametric study like the change of total flow rate study are presented elsewhere [16,38].
3. Research method 3.2. Numerical model and solution procedures 3.1. Experimental facility In order to figure out the feature of the thermal decomposition of R-134a as a function of reactor wall temperature and oxygen flow rate, a series of experiments have been made using tubular reactor made of Hastelloy material as shown in Fig. 1. In order to supply the activation energy of R-134a reaction, the tubular reactor wall temperature is heated to a desired temperature. In specific, the furnace diameter and the length of reactor are 22 mm and 600 mm respectively and only 500 mm length was heated from 100 to 900 °C with 100 °C interval, while the flow rate of R-134, water vapor and N2 are fixed with 100 g/h, 17.7 g/h and 100 g/h respectively. The flow rate of oxygen and water vapor are supplied
Numerical calculation has been made by a commercial code, STAR-CCM+(Ver.9.06), which can deal with the pre-process, solver and post-process using the integrated package CAE (computer aided engineering) software in one working environment, and can solve the physical phenomenon in various fluid flow and heat transfer problems associated with chemical reaction. The same input condition was employed with the experimental condition with the flow rate of R-134a 100 g/h together with the O2 47.1 g/ h, H2O 17.7 g/h and carrier gas, N2 100 g/h via the gas mixing chamber. For completeness, the basic gas phase conservation equations for mass, momentum, energy, turbulent quantities and species concentration can be expressed as in an Eulerian framework,
Fig. 1. Detailed specification of tube reactor and boundary condition.
Table 1 General dependent variable / and diffusion coefficient C/ expression for axi-symmetric cylindrical coordinate system for the general governing Eq. (22). C2 H2 F4 þ 32 O2 þ H2 O þ 3:76N2 ! 4HF þ 2CO2 þ 3:76N2 ð22Þ. Variables
/
C/
Axial momentum
u
leff
Radial momentum
v
leff
Tangential momentum
w
leff
Kinetic energy
k
Kinetic energy dissipation rate
e
leff rk leff rs leff rh
S/
@ @x @ @x
@p 1 @ @v leff @u @x þ r @r leff r @x @x
qw @p 1 @ @v v leff @u @r þ r @r leff r @r 2leff r2 þ x @r
l
eff
r2
þ qrv þ
Gk1 qe
e j ðC 1 C k1 C 2 qeÞ
_ fu Hfu w
Specific enthalpy
h
Fuel mass fraction Oxygen mass fraction
mfu mO2
Cfu CO2
_ fu w
HF mass fraction
mHF
CHF
4
_ fu w M fu
CO2 mass fraction
mCO2
CCO2
2 Mfufu MCO2
H2O mass fraction
mH2 O
CH2 O
Mfufu MH2O
Gk1 ¼ 2leff
h @u 2 @x
þ
@ v 2 @r
h 2 i @ w2 @u @ v 2 i 2 þ vr þ r @r þ @r þ @x þ leff @w @x r
C1 = 1.44, C2 = 1.92, Cl = 0.92, rk = 0.9, re = 1.22
1 @ leff r @r
_ w
1:5 Mfufu M O2 MHF
_ w
_ w
w
2
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M. Shin et al. / Applied Thermal Engineering 125 (2017) 20–28 !
where / represents the general dependent variable, and q and u are density and velocity vector, respectively. Further, C/ and S/ are the effective diffusion coefficients and the source term respectively for /. ! @ðq/Þ þ r ðq u /Þ ¼ r ðC/ r/Þ þ S/ @t
ð20Þ
The popular approach is the use of the two equation (k–e) model by Launder and Spalding, where a Prandtl-Kolmogorov relationship is used to correlate the turbulence viscosity, lt to the turbulent kinetic energy (k) and its dissipation rate (e) [39].
lt ¼ Cl qk2 =e
ð21Þ
As mentioned above, for the prediction of the non-equilibrium thermal destruction of HFC-134a, a hybrid type eddy breakup model is developed by the consideration of the relative rate of the turbulent mixing and chemical reaction together. To this end, empirical constants in the chemical reaction rate are determined by the trial and error method based on the experimental data in this tubular
Fig. 2. Comparison of thermal decomposition of R-134a at the reactor exit between numerical calculation and experiment as a function of wall temperature.
reactor. Further the hybrid type EBU model is employed for the prediction of the non-equilibrium turbulent reaction in this tubular reactor. The diffusion coefficient C/ and source term S/ in Eq. (20) are presented for axi-symmetric cylindrical coordinate system in Table 1 for the dependent variables such as u, v, k, e, h, mfu ; mO2 ; mHF ; mCO2 ; mH2 O respectively. The dependent variable u, v, h, mfu ; mO2 , mHF ; mCO2 , mH2 O are denoted axial and radial velocities, enthalpy, mass fractions of fuel, oxygen, hydrogen fluoride, carbon dioxide and water, respectively. In this calculation, a suitable radiation model was tested but no visible change was observed, as expected, probably due to the low emissivity by the small mean beam length caused by the small tube diameter. Fig. 2 represents a boundary condition of tube reactor for numerical calculation. 4. Results and discussion Experimental data of thermal decomposition at reactor exit as a function of temperature and oxygen concentration are presented in Tables 2 and 3, respectively. In Table 2, the fraction of the
Fig. 3. Comparison of thermal decomposition of R-134a between numerical calculations by two EBU models and experiment as a function of oxygen mole flow rate.
Table 2 Experimental data of thermal decomposition of R-134a as a function of temperature with fixed amounts of O2 and H2O. (*Carrier gas N2 100 g/h). Expt no.
Reactor wall temperature (°C)
C2H2F4 (g/h)
O2 (g/h)
H2O (g/h)
Mole ratio (C2H2F4:O2:H2O)
Decomposition percent (%) (experiment)
Decomposition percent (%) (calculation)
1 2 3 4 5 6 7 8
100 300 500 600 700 800 850 900
100 100 100 100 100 100 100 100
47.1 47.1 47.1 47.1 47.1 47.1 47.1 47.1
17.7 17.7 17.7 17.7 17.7 17.7 17.7 17.7
1:1.5:1 1:1.5:1 1:1.5:1 1:1.5:1 1:1.5:1 1:1.5:1 1:1.5:1 1:1.5:1
0 0 0 1.1 17.7 79.6 100 100
0 0 0 5.2 20.7 80.4 100 100
Table 3 Experimental and numerical data of the thermal decomposition of R-134a as a function of oxygen flow using two EBU models (*Carrier gas N2 100 g/h and 900 °C wall temperature and the stoichiometric amount of H2O.) Expt no.
Reactor wall temperature (°C)
C2H2F4 (g/h)
O2 (g/h)
H2O (g/h)
Mole ratio (C2H2F4:O2:H2O)
Experimental decomposition percent (%)
Numerical decomposition percent (%) by fast chemistry EBU model
Numerical Decomposition percent (%) by hybrid model
8 9 10 11 12 13 14
900 900 900 900 900 900 900
100 100 100 100 100 100 100
47.1 0 9.4 15.7 21.9 28.3 37.7
17.7 17.7 17.7 17.7 17.7 17.7 17.7
1:1.5:1 1:0.0:1 1:0.3:1 1:0.5:1 1:0.7:1 1:0.9:1 1:1.2:1
100 25 50.8 – – 94 100
100 31.8 53.4 67.4 77.6 86.8 94.6
100 22.7 57.1 75.7 86.2 94.7 100
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M. Shin et al. / Applied Thermal Engineering 125 (2017) 20–28
HFC-134a thermal decomposition is presented with the wall temperature increase from 100 °C to 900 °C for the fixed value of the stoichiometric amount of oxygen and water vapor. In Table 3, the thermal decomposition data are presented as a function of oxygen concentration with the fixed value of wall temperature and 900 °C and stoichiometric amount of H2O concentration. As shown in Table 2, no thermal decomposition of HFC 134a occurs up to the wall temperature of 500 °C and only very small amount of decomposition, that is, 1.1% occurs at temperature 600 °C even if there exist stoichiometric amount of oxygen and water vapor. But the fraction of thermal decomposition increases 17.7, 79.6 and 100% as wall temperature increases 700, 800, and 850 °C, respectively. This data suggests that at least more than 600 °C is necessary in order to react HFC-134a with O2 and water vapor. Using this
decomposition data at the reactor exit, empirical constants of pre-exponential factor (A), empirical constant B and activation Ea
energy (Ea) in the chemical reaction rate, that is, k ¼ AT B eðRT Þ was determined by trial and error method in order to fit the destruction fraction of HFC-134a in Table 2 using the hybrid type EBU model for the development of the non-equilibrium turbulent reaction model of HFC-134a incineration. The value of A, b and Ea obtained are A = 1.0, B = 0.38, Ea = 8.8 107 J/kmol, respectively. The calculated thermal decomposition data are well compared with experimental data and presented in Fig. 2 and Table 2, respectively. Further, using the hybrid EBU model with empirical constants determined above, the thermal destruction data as in Table 3 are predicted as a function of oxygen concentration with fixed wall
(a) Velocity vector( inlet velocity,0.126 m/s)
(b) Temperature distribution
(c) Mass fraction of HFC-134a Fig. 4. Calculation results for the wall temperature of 900 °C and stoichiometric condition of oxygen and water vapor by conventional EBU model for the Exp. No. 8 in Table 2 or 3.
(a) Velocity vector( inlet velocity,0.126 m/s)
(b) Temperature distribution
(c) Mass fraction of HFC-134a Fig. 5. Calculation results for the wall temperature of 900 °C and stoichiometric condition of oxygen and water vapor by the non-equilibrium EBU model for the Exp. No. 8 in Table 2 or 3.
M. Shin et al. / Applied Thermal Engineering 125 (2017) 20–28
27
Fig. 6. Comparison of numerical calculations by two EBU models along the axial distance of tube reactor.
temperature 900 °C. In this calculation, the fast chemistry EBU model used in Shin et al. [16] is also employed for the comparison purpose. Comparison between calculations by two EBU models and experiment has been made and presented at Table 3 and in Fig. 3 for the thermal decomposition of R-134a as a function of O2 concentration. In general, the prediction by hybrid-type EBU model agrees better than the conventional fast chemistry model without any non-equilibrium incorporation. Since the comparison has been made only for the concentration of HFC-134a at the exit, the calculated profiles are presented in detail in order to see the difference of two models. In Figs. 4 and 5, the detailed profiles of HFC-134a concentration, temperature and velocity field are presented respectively. Even if the overall thermal decomposition fraction by two models look similar at the exit, the profiles of temperature and HFC-134a concentration differ significantly. As shown in Fig. 5, the reaction of HFC-134a occurs very slowly at low temperature entrance region due the non-equilibrium effect. Further, detailed profile comparisons of temperature and HFC-134a have been made along the reactor centerline axis using two models as shown in Fig. 6. This clearly shows the non-equilibrium effect of turbulence reaction model plays a significant role especially in the significant region of the non-equilibrium chemistry effect especially in low flame temperature region.
5. Conclusion Considering the global warming potential of HFC-134a (C2H2F4) with the substantial generation of this refrigerant as waste material in various industrial sectors, the development of proper thermal destruction method of HFC-134a is of great practical significance especially in the generation of fluorine material like HF and CaF2. Experiment and numerical calculation have made for a tubular type in order to figure out the important feature of the thermal decomposition of HFC-134a as a function of temperature and reactants such as oxygen and water vapor. Detailed review and discussions are made for the modeling and nonequilibrium effect in the incineration process of various halogenated hydrocarbons using a couple of modified versions of eddy breakup (EBU) model developed from the original EBU model. Employing this non-equilibrium hybrid EBU model, the calculated overall destruction rate of HFC-134a was in good agreement with the experimental data obtained for the same reactor. But the detailed HFC-134a concentration profiles between the fast chemistry and non-equilibrium EBU models show a significant difference even if the calculated destruction rates of HFC-134a appear similar at the exit. In near future, numerical calculation will be
made for the practical incinerator of HFC-134a in a practical incinerator using the model developed in this study. Acknowledgement This project is supported by the ‘‘R&D Center for reduction of Non-CO2 Greenhouse gases (2013001690006)” funded by Korea Ministry of Environment (MOE) as ‘‘Global Top Environment R&D Program”. References [1] China Fluoro Technology Co., CDM Monitoring Report, UNFCCC (United Nations Framework Convention on Climate Change), CDM Reference Number 1194, 2007. [2] D.S. Jang, M.S. Shin, Y.G. Lee, Y.J. Kim, On a thermal destruction of HFC-134a and HFC-23 by CH4 and water electrolysis gas in a CDM incinerator, in: The 2nd 3R International Scientific Conference on Material Cycles and Waste Management, 2015, pp. 205–209. [3] US Environmental Protection Agency, Complying with the Section 608 Refrigerant Recycling Rule, 2011.
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