Coke Deposition Influence Based on a Run Length ... - ACS Publications

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Coke Deposition Influence Based on a Run Length Simulation of a 1,2-Dichloroethane Cracker Chaochun Li, Guihua Hu, Weimin Zhong, Wangli He, Wenli Du,* and Feng Qian* Key Laboratory of Advanced Control and Optimization for Chemical Processes, Ministry of Education, East China University of Science and Technology, Shanghai 200237, China ABSTRACT: A full cycle of an industrial ethylene dichloride cracker is simulated. Given the intense heat coupling between the furnace and the reactor, the cracker is divided into two parts: the furnace model and the reactor model, with heat flux and flue gas temperature profiles connecting the two models. A radical mechanism with coke formation is adopted to describe the EDC cracking reactions with 24 reaction equations and 31 components. In the full cycle simulation, two important aspects, namely, CCl4 concentration and fuel gas allocation, are investigated to understand the overall benefits of the whole operation cycle. Addition of the promoter CCl4 to EDC raw material can improve EDC conversion. However, this process aggravates the coking reaction, which causes the sharp deterioration of the cracking performance and the shortening of the running cycle. On the other hand, the fuel gas allocation factor facilitates analysis of the fuel gas allocation strategies. Increasing the fuel gas amount at the furnace bottom can effectively improve the heat transfer efficiency of the EDC cracker. In particular, this process enhances heat transfer at the end of the tubular reactor, which improves the EDC conversion. However, coke deposition greatly shortens the run cycle. A comprehensive analysis shows that the concentration of the CCl4 promoter should be controlled at 100 ppm wt % and the fuel gas allocation factor should be maintained at 0.36 to guarantee the overall economic benefits of the EDC cracker in the full operation cycle.

1. INTRODUCTION Poly vinyl chloride (PVC), the second most used plastic worldwide, is produced by the polymerization of vinyl chloride monomers (VCMs). The pyrolysis of 1,2-dichloroethane, also named ethylene dichloride (EDC), is the only commercial route to form VCM. Given the high energy consumption of the EDC cracking process, the EDC furnace is considered as the heart of the entire VCM manufacturing process. A typical EDC furnace generally includes the convection section and the radiation section. The EDC feedstock is preheated in the convection section and then introduced to a long tubular reactor that lines the center of the radiation section for endothermic cracking reactions. Fuel gas and air are injected into the furnace through burners to combust and raise the process gas temperature and support the endothermic reaction. The reactions in the reactor are strongly affected by the processes in the furnace and vice versa.1 The structure diagram of an EDC cracker is illustrated in Figure 1. In principle, the complex thermal cracking of EDC is considered to proceed via free-radical reactions. Rigorous reaction mechanisms have been studied and improved several times by various researchers.2−9 Ranzi et al.6 introduced a reaction kinetic model with more than 200 elementary reactions and more than 40 molecular species and radicals involved in the EDC pyrolysis reaction network. Borsa et al.7 developed the most detailed chemical kinetics model for EDC pyrolysis. As many as 135 compounds and radical species and more than 800 reactions have been reported for comprehensive calculations. Choi et al.8 established a mechanism that involves 108 reversible elementary reactions and 47 molecular/radical species. The addition of CCl4 was first investigated as a process by Choi et al.9 Schirmeister et al.10 simplified the foregoing EDC pyrolysis mechanism in the view of data accuracy and © 2013 American Chemical Society

expenditure optimization for model adjustment. A total of 31 reactions, 18 compounds, and 8 radial species were used to describe all relevant products, intermediates, and byproducts. EDC thermal cracking always accompanies coke deposition. Several published reports analyzed coke classifications and locations using electronic analysis equipment in situ or in laboratories.11−13 Coke formation is known to cause several undesirable consequences on the EDC pyrolysis process. First, considering that the coke layer is a poor heat conductor, coke formation leads to decreased furnace thermal efficiency. Furthermore, increased coke thickness results in increased skin temperature of the reactor, which is the main limitation of coil metallurgy. Moreover, the increase in coke layer thickness reduces the cross section of the reactor tube. With a constant EDC feed rate, the pressure drop along the tubes increases and the residence time decreases. Given that the process gas outlet pressure is normally imposed according to the downstream operation condition, the feed inlet pressure increases during the run length of the furnace. Reduced EDC cracking conversion and increased maintenance and utility costs are the other results of the coke formation phenomenon. These problems force plant operators to shut down the unit for decoking process. A typical run length of a commercial EDC cracker is 1 year to 2 years, which depends on the type of reactor, feedstock, and operating conditions. The decoking process is carried out by burning the coke using a mixture of air and steam. Consequently, the coil lifetime is gradually shortened due to Received: Revised: Accepted: Published: 17501

April 24, 2013 July 19, 2013 October 23, 2013 October 23, 2013 dx.doi.org/10.1021/ie401265f | Ind. Eng. Chem. Res. 2013, 52, 17501−17516

Industrial & Engineering Chemistry Research

Article

Figure 1. Ethylene dichloride cracker structure diagram.

balancing the EDC cracking performance indices with the coke formation. The run length simulation of the thermal-coupled EDC cracker is presented in three parts, namely, the reactor model, the furnace model, and the kinetic model, with coke formation. Coke layer thickness, external/internal tube skin temperature distributions, EDC cracking conversion, and other important EDC cracking performance indices are precisely predicted with the run time. The external skin temperature limitation caused by coil metallurgy is used to predict the run length of the EDC cracker. As soon as the maximal external skin temperature reaches the limit, the EDC cracker should be shut down for decoking. Meanwhile, based on the EDC cracker simulation, two aspects are considered to reduce coke formation, cut down economic cost, improve EDC cracker operating rate, and increase VCM production capacity. Primarily, the concentration of CCl4, which is taken as the EDC cracker promoter, is investigated. An approximate CCl4 concentration is suggested to promote EDC pyrolysis, reduce the energy consumption of the unit VCM yield, and guarantee no excessive production of coke. Furthermore, the fuel gas allocation is also discussed with the definition of the fuel gas allocation factor. The distribution of the external tube skin temperature and EDC performance indices are analyzed by adjusting the fuel gas allocation factor. A suitable fuel gas allocation strategy is presented to improve the EDC cracking performance indices and avoid excessive loss of the run time. This article is presented in the following arrangement. The second section describes the mathematical model of the EDC cracker, which involves the EDC cracking free-radical model with coke formation, reactor model, and furnace model. The numerical iterative method of the thermal-coupled furnace and reactor model is also presented. The third section describes the simulation based on the run length of an industrial MITSUI EDC cracker. The fourth section explores the influence of CCl4 concentration on the EDC cracking process and coke formation. The fifth section discusses the impact of different fuel gas allocation factors on the cracking performance and operation cycle. A conclusion is drawn in the sixth section.

constant thermal cycling. The qualitative and quantitative analysis of coke influence on EDC cracker based on the run length simulation is very critical. Predicting the run time and extending the operation cycle of an EDC cracker by minimizing the amount of coke deposition is highly necessary. Several run length simulation research studies of hydrocarbon crackers have been performed and presented in the literature.14−18 However, only a few research studies on the run length simulation of EDC cracker has been reported. In the EDC cracking process, the coke rate is mainly determined by two operational factors, namely, the purity of the raw material and the heat flux distribution along the reactor. A typical EDC feed purity is required to be greater than 99%. However, the rest of the minor components still pose an important impact in the EDC pyrolysis process, particularly the substances rich in Cl element. The characteristics of Clcatalyzed reaction mechanisms allow Cl suppliers to play key roles in promoting the reaction conversion. Huybrechts et al.4 discussed the different situations of EDC pyrolysis in the absence and presence of added Cl2. Ranzi et al.6 and Choi et al.8 also analyzed the effects of the addition of different concentrations of CCl4 as promoter. Although the promoters accelerate the EDC cracking rate, improve EDC conversion, and reduce fuel gas consumption of unit VCM production, they also aggravate undesirable coke formation. Thus, selecting a suitable promoter concentration is necessary in optimizing the EDC cracking efficiency and economy while minimizing coke formation and extending the run length of the EDC cracker. Furthermore, the endothermic cracking process in the reactor tube is strongly affected by the processes in the furnace and vice versa. The heat flux profile along the reactor tubes forms the connection between the process and the flue gases. The heat flux profiles generated in the furnace are transferred to the process gas inside the reactor tubes, which raises the process gas temperature and supports endothermic cracking reactions along the reactor tubes. The heat is generated by the combustion of the fuel gas, which is injected into the furnace from the burner lines at different positions. Consequently, the flue gas temperature distribution in the furnace and the heat flux along the reactor is directly determined by the fuel gas allocation in the different burner lines. The selection of an approximate fuel gas allocation strategy is another core focus in 17502

dx.doi.org/10.1021/ie401265f | Ind. Eng. Chem. Res. 2013, 52, 17501−17516


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2. MATHEMATICAL MODEL 2.1. Kinetic Model. The radical chain mechanism of thermal cracking is well-known in principle. Experimental studies and rigorous mathematical models of the complex EDC cracking reaction can be found in the literature.2−10 In this research, the simplified radical chain mechanism proposed by Schirmeister10 is adopted to describe the EDC cracking process, which results in 31 reactions that describes all 24 relevant products, intermediates, and byproducts. Table 1 lists all 16 “stable” substances and 8 radicals and their brief descriptions.

stream. The simulation of the reactor integrates a set of continuity equations for the process gas species, along with the energy and momentum equations. The following assumptions are made to simplify the mathematical model: 1. one dimensional plug flow and laminar regime 2. ideal gas behavior 3. negligible radial gradient and axial dispersion 4. no hydrodynamic or thermal entrance region effects 5. quasi steady-state in coke deposition model With the above assumptions, the governing equations in steady state proposed by Plehiers30 are derived as follows: dFj

Table 1. Compounds, Radical Species, And Short Forms in the Kinetic Reactions no.

compound

dz

short form

∑ Fjcpj

Monomolecular Species 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 1 2 3 4 5 6 7 8

1,2-dichoroethane vinylchoride hydrogen chloride trichloromethane tetrachloromethane ethylchloride 1,1-dichloroethane 1,1,2-trichlorothane 1,1,1,2-/1,1,2,2-tetrachloroethane 1,1-/cis-/trans-dichloroethylene trichloroethylene 1-/2-chloroprene acetylene benzene 3,4-dichlorobutene soot/coke Radical Species Cl* CH2Cl-CH2*/CH3−CHCl* CH2Cl−CHCl* CHCl2−CH2* CHClCH*/CH2CCl* CH2Cl−CCl2*/CHCl2−CHCl* CHCl2−CCl2*/CCl3−CHCl* CCl3*

= (∑ sijrri)

EDC VCM HCl CHCl3 CCl4 EC 1,1 1,1,2 1,1,1,2 Di Tri CP C2H2 C6H6 C4H6Cl2 C

j

i

πd t 2 4

πd t 2 dT ≡ Q (z )π d t + dz 4

(1)

∑ rri(−ΔH )i i

(2)

⎛ 1 Pt ⎞ dpt ⎞ d⎛ 1 ⎞ 1 ⎛⎜ 1 dT − = + Fr ⎟ ⎜ ⎟ ⎟+ ⎜ 2 ⎠ ⎝ dz ⎝ M m ⎠ M m T dz ηG RT ⎠ dz ⎝ M mPt (3)

The friction factor for the straight parts of the cracking coils and tube bends are given in eqs 4 and 5, respectively: Fr = 0.092

ζ Re−0.2 + πR b dt

⎛ dt ⎞ Λ ⎞⎛ ⎟⎜0.051 + 0.19 ζ = ⎜0.7 + 0.35 ⎟ ⎝ 90° ⎠⎝ Rb ⎠

(4)

(5)

where Rb and ζ represent the tube bend radius and bend angle, respectively. According to the coking model shown in Table 2, acetylene is assumed to be the only coke precursor as the description of reaction 31. To predict coke deposition, the rate of coke formation is expressed as

R1 R2 R3 R4 R5 R6 R7 R8

⎛E ⎞ rc = A31 exp⎜ 31 ⎟ × CC2H2 × C R1 ⎝ RT ⎠

(6)

Coke thickness is related to the time and position of the reactor and determined by the following equation by Sundaram and Froment.31

Table 2 lists the 31 kinetic reactions and kinetic data. The presented radical mechanism has the same steps as a typical radical mechanism, which includes chain initiation, chain propagation, and chain termination. Some simplifications are done in the present kinetic model. As Cl* play a very active role in the radical mechanism of EDC pyrolysis, CCl4 is emphasized and considered as the promoter to provide abundant Cl* radicals to be propitious for EDC cracking. Furthermore, a number of coke precursors have been widely studied and reviewed in chlorinated hydrocarbon pyrolysis or nonchlorinated hydrocarbon pyrolysis,9−11,19−29 in which acetylene is always considered as an important possible coke precursor in the coke formation. However, in the present study, the coke formation mechanism is simplified as the Schirmeister work.10 Acetylene is taken as the only precursor for carbon formation without the consideration of the complex synthetic pathways. 2.2. Reactor Model. In the present study, a onedimensional plug flow reactor model is assumed based on a high Reynolds number and a low viscosity of the reaction side

⎡ aM r ⎤ dδ(z) = (d i − 2δ(z))⎢ c c ⎥ ⎢⎣ 4ρc ⎥⎦ dt

(7)

Given that the coke formation is slow, quasi-steady state conditions are assumed such that the coking rate is constant during the sampling interval. The time differential term is removed from the coke thickness eq 7 with the definition of thickness increment Δδc(z). Δδ(z) =

aMcrc Δt ρc

(8)

Consequently, the flow diameter of the reactor in the numerical iterative algorithm could be updated as follow: d t_new(z) = d t_old(z) − 2Δδc(z)

(9)

From the energy balance in eq 2 of the governing equations, the EDC cracking process is shown to be supported by the heat 17503



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Table 2. Kinetic Reactions and Kinetic Data no.

reactions

frequency factor [(cm3/mol)n−1s−1]

n

activation energy Ea (kJ/mol)

Radical reaction chain initialization 1 EDC → R1 + R2 5.9 × 1015 1 342 2 CCl4 → R1 + R8 2.2 × 1012 1 230 Radical reaction chain propagation: (a) H* abstraction (3−9,11, and 19−23), (b) Cl* formation (15−19), (c) double bond formation (10, 12−14), (d) Cl* abstraction from CCl4 (24−26) 3 EDC + R1 → HCl + R3 1.3 × 1013 2 7 4 EDC + R5 → VCM + R3 1.2 × 1013 2 34 2 42 5 EDC + R2 → EC + R3 1.0 × 1012 6 EDC + R4 → 1,1 + R3 5.0 × 1011 2 45 7 EDC + R6 → 1,1,2 + R3 2.0 × 1011 2 48 8 EDC + R7 → 1,1,1,2 + R3 1.0 × 1011 2 56 9 EDC + R8 → CHCl3 + R3 1.0 × 1012 2 63 2 0 10 VCM + R1 → R4 9.1 × 1010 11 VCM + R1 → HCl + R5 1.2 × 1014 2 56 12 VCM + R5 → CP + R1 5.0 × 1011 2 31 13 VCM + R4 → C4H6H2 + R1 2.0 × 1010 2 30 2 61 14 VCM + R2 → EC + R5 3.0 × 1011 15 R3 ↔VCM + R1 2.1 × 1014 1 84 16 R5 ↔C2H2 + R1 5.0 × 1014 1 90 1 70 17 R6 ↔Di + R1 2.0 × 1013 18 R7 ↔Tri + R1 2.5 × 1013 1 70 19 2C2H2 + R5 → C6H6 + R1 1.0 × 1014 2 20 20 EC + R1 → HCl + R2 1.7 × 1013 2 4 2 6 21 1,1 + R1 → HCl + R4 1.2 × 1013 22 1,1,2 + R1 → HCl + R6 1.7 × 1013 2 15 23 1,1,1,2 + R1 → HCl + R7 1.7 × 1013 2 17 24 CHCl3 + R1 → HCl + R8 1.6 × 1013 2 14 2 33 5.0 × 1011 25 CCl4 + R5 → Di + R8 26 CCl4 + R4 → 1,1,2 + R8 2 33 1.0 × 1012 27 CCl4 + R5 → 1,1,1,2 + R8 5.0 × 1011 2 33 Radical reaction chain termination 28 R2 + R1 → VCM + HCl 1.0 × 1013 2 13 29 R3 + R1 → Di + HCl 1.0 × 1013 2 12 30 R6 + R8 → Di + CCl4 1.0 × 1013 2 13 1.6 × 1014 2 70 31 C2H2 + 2R1 → 2C + 2HCl

transferred from the reactor tube. Tube thickness, coke

from this setup, two fouling layers are introduced to improve the flexibility and accuracy of the reactor tube heat transfer calculation and are taken as a useful supplement of the simplified coke model. One layer, which is attached on the coke layer, is believed to be the film of tar liquefied from process gas. The other layer is the fouling layer caused by ash-accumulation on the external reactor skin. The heat resistances are estimated at 6.8 × 10−5 m2·K·W−1, 4.3 × 10−5 m2·K·W−1, respectively, without consideration of the layer thickness. The heat flux profile transferred from the local position reactor is given by eq 10:

thickness, and the gaseous phase are the main factors that account for the thermal resistance, as shown in Figure 2. Apart

Q t(z) = K (Two(z) − T (z))

(10)

with the total thermal conductivity calculated as follows: d d d d d 1 1 do = + R i o + o ln i + t ln o + R o K α dt dt 2λc dt 2λ w di (11)

2.3. Furnace Model. As EDC cracking reactions mainly occur in the radiation section of the furnace after sufficient preheat of the EDC feedstock in the convection section, only the radiation section is considered in the furnace model. On the basis of the clear temperature gradient along the height inside the EDC cracker furnace, the radiation section of the furnace is

Figure 2. Different layers in the ethylene dichloride cracking reactor tubular. 17504



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composition is listed in Table 3, the C1, C2, and H2 contents in the fuel gas are relatively small and not explicitly accounted, which slightly affect the fuel gas combustion. In the simulation, these contents in the fuel gas are assumed to be replaced by H2. Fuel gas is preheated to 80 °C first and then injected to the furnace with natural air inlet around the fuel gas nozzles. The temperature of the air is supposed to be 20 °C. According to low heat value of each component of the fuel gas, the total low heat value of the fuel gas Qsl could be calculated by the following equation:

divided into several zones, as shown in Figure 3. Onedimensional Lobo−Evans method is used to account for the furnace model.

Q sl =

∑ Yfi × Q fli i

(12)

where theYfi is the mass friction of fuel gas component i, Qfli is the low heat value of fuel gas component i. Heat is generated in the each furnace partition could be calculated with the assumption of complete combustion in its local partition. Q si = GsiQ sl

Figure 3. Structure of the radiation section in the furnace.

where Qsi is the heat released by the fuel gas in the ith zone, Gsi is the fuel gas mass flow sprayed in the ith zone. Heat transferred from the high-temperature flue gas to the reactor tube consists of two parts, namely, radiation and convective heat transfers, as defined below.

Heat is generated in the furnace through the combustion of the fuel gas. Fuel gas allocated in each furnace partition is assumed to be totally combusted in to CO2 and H2O directly in one-step reaction mechanism as the following expression ⎛ y⎞ y Cx Hy + ⎜x + ⎟O2 → xCO2 + H 2O ⎝ ⎠ 4 2

Q ri = σaDAcpi Fi(Tgmi 4 − Twi 4) + hRcARi (Tgmi 4 − Twi 4) (14)

Details of the firing conditions are listed in Table 3. To guarantee the total combustion of fuel gas, the oxygen excess coefficient is set at 3.0% (vol). Theoretical air consumption of the fuel for combustion could be calculated based on the fuel gas composition. Considering the fuel used in the trial whose

Each furnace zone should meet the energy balance principle, with five parts in consideration: heat released by the fuel gas Qsi, net radiation heat brought in/out from the axial neighboring zone, heat absorbed by the reactor tubes Qri, net heat of flue gas enthalpy change ΔHi, and furnace wall heat dissipation Qli. Therefore, the energy balance equation is defined as follows:

Table 3. Furnace Dimension and Operating Conditions

⎧ 4 4 ⎪ Q si + σA 0[Tgm2 − Tgm1 ] − Q r1 − ΔH1 − Q l1 ⎪ =0 ⎪ ⎪ Q s2 + σA 0[Tgm14 − Tgm2 4] + σA 0[Tgm3 4 − Tgm2 4] ⎪ ⎪ − Q r 2 − ΔH1 − Q 12 = 0 ⎪ ⎪⋮ ⎪ 4 4 ⎪ Q si + σA 0[Tgm(i − 1) − Tgmi ] 4 4 ⎪ ⎪ + σA 0[Tgm(i + 1) − Tgmi ] − Q r1 − ΔHi − Q li ⎨ ⎪ =0 ⎪⋮ ⎪ ⎪Q + σA 0[Tgm(N − 2)4 − TN − 14] ⎪ s(N − 1) ⎪ + + σA 0[Tgm(N )4 − TN − 14] − Q r(N − 1) − ΔHN − 1 ⎪ ⎪ − Q l(N − 1) = 0 ⎪ ⎪ Q sN + σA 0[Tgm(N − 1)4 − Tgm(N )4] − Q rN − ΔHN ⎪ ⎪ ⎩ − Q ln = 0

Furnace Segment length (m) (x-direction) width (m) (y-direction) height (m) (z-direction) number of wall burners Firing Conditions

(13)

20.898 1.900 6.700 136

fuel gas flow rate (kg/h) oxygen excess (vol %) heat loss of the furnace wall (%) temperature of fuel gas (°C) temperature of air (°C) C4H8 C4H6 C4H10 C3H8 H2, C1, C2 Reactor Tube

999.3 3% 3.0% 80 20 76.0% 20.0% 2.6% 0.7% 0.7%

number of reactor row number of stright tubes (per row) number of bends (per row) inlet tube diameter (10−3 m) outlet tube diameter (10−3 m) thickness of tube (10−3 m) EDC Feed rate (t/h) coil inlet temperature (K) coil outlet pressure (kPa)

2 20 19 101.3 114.3 6.5 42 533.15 2400

(15)

2.4. Physical Properties. The process gases inside the reactor tube are assumed as ideal. Thus, the density of the mixture can be calculated by the ideal gas law: p ρ= Y RT ∑ Mi (16) i

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Figure 4. Flow diagram of the iterative solution algorithm of thermal coupled EDC cracker.

where Ai, Bi, Ci, and Di are the coefficients of heat capacity of pure species i. For the thermal conductivity of pure species, the kinetic theory is used:

The viscosity of mixture is given by μ=

∑ i

Xiμi ∑j Xiϕij

(17)

The viscosity of the pure species i being calculated as μi = 2.67 × 10−6

λi =

Mi T σi Ωμi

[1 + (μi /μj )1/2 (Mi /Mj)1/4 ]2 {8[1 + (Mi /Mj)]}1/2

λ=

∑ YC i pi i

X i λi ∑j Xjϕij

(23)

For the calculation of the physical properties of the flue gas, only the heat transferred model is considered in the Lobo− Evans method. The used method for heat capacity is similar to that of the process gas. 2.5. Coupled Numerical Simulation between the Furnace and the Reactor. The simulation of the thermal coupled EDC cracker is achieved using Matlab(2009b). Heat is strongly coupled between the furnace and the reactor. Hence, the heat flux transferred from the furnace to the reactor must be balanced with that absorbed by the process gas along the

(20)

where Cpi is the heat capacity of pure species i and defined by a polynomial function of temperature: Cpi = Ai + Bi T + CiT 2 + DiT 3

∑ i

(19)

The heat capacity of the mixture is computed from a mass fraction average of the heat capacities of pure species as

Cp =

(22)

The thermal conductivity of the mixture can then be computed by

(18)

In eq 14, the interaction parameter ϕij is calculated by ϕij =

15 R ⎡ 4 CpiMi 1⎤ μi ⎢ + ⎥ 4 Mi ⎣ 15 R 3⎦

(21) 17506



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reactor length. In this work, the coupled numerical solution of the one-dimensional Lobo−Evans method is adopted, and flue gas temperature and heat flux profiles are considered the core variables. The boundary conditions of the inlets in the furnace and the reactor are determined according to Table 3. Details of the flow diagram of numeric calculation are illustrated in Figure 4 as the following steps: 1: Some necessary specifications of the furnace and reactor (such as coil geometry, feedstock condition and so on) are initialized. 2: The flue gas temperature profiles in the furnace are estimated and passed to the reactor model and coil inlet pressure (CIP) in the reactor is also estimated. 3: Heat flux profiles along the reactor length are calculated in the reactor model through the solution of governing equations including the mass, momentum and energy balance equations, and then are passed to the furnace model. 4: Flue gas temperature profiles along the furnace height are calculated through the energy balance solution in the furnace. 5: The convergence of the flue gas temperature profiles is judged by comparing the new-calculated flue gas temperature profiles with flue gas temperature profiles calculated last time. If the flue gas temperature profiles are converged, turn to step 6. If not, then the flue gas temperature profiles are assigned to the reactor model again. The circulatory iterative calculation turns in the step 3. 6: The convergence of coil outlet pressure (COP) in the reactor is examined by comparing the new-calculated COP with the set-point. If the COP meets the computation accuracy, then turns to step 7. If not, the CIP is re-estimated and the circulatory iterative calculation turns to step 3. 7: Maximum temperature of external reactor skin is checked. If it reaches the withstanding limit, the whole calculation is terminated. If not, the flow diameter profiles of reactor are updated. The circulatory iterative calculation goes to step 2, and calculation of a new sampling-time interval of the whole run-length of EDC cracker starts. The integration of the mass balances in the reaction model is described using ordinary differential equations (ODEs). Since 24 species and 31 reactions exist in the reactor, a stiff problem is required to solve the set of equations. One of the built-in ODE solvers in the Matlab, named ODE23s, which is based on a modified rosenbrock formula of order 2, is used to solve the ODE stiffness problem. The integration step size of the ODEs is automatically controlled and the absolute error tolerance of that is set at 1 × 10−6. In the numeric iterative simulation, convergence criteria of coil outlet pressure (COP) are set at 0.1 kPa, while the Convergence criteria of the flue gas temperature is set at 0.1K.

Table 4. Comparison of Simulation Results and Industrial Data for Clean Tube Stage item temperature of fuel gas at outlet (K) temperature of process gas at outlet (K) coil inlet pressure (kPa) coil pressure drop (kPa) EDC molar flow rate at reactor outlet (kmol/h) VCM molar flow rate at reactor outlet (kmol/h) HCl molar flow rate at reactor outlet (kmol/h) conversion selectivity fuel gas consumption per VCM production (kg/t)

industrial data

simulation data

relative error

900.0 763.0

940.0 764.7

4.44% 0.22%

3000.0 6000.0 95.462

3015.4 6015.4 90.4729

0.51% 2.57% 5.23%

115.115

120.006

4.25%

117.075

120.792

3.17%

0.555 0.978 69.45

0.575 0.986 66.6166

3.60% 0.82% 4.01%

the industrial data at the clean tube stage. It can be seen that the model is acceptable at a certain accuracy level with the relative error below 5.3%. It is believed that two factors are responsible for the model errors in the present work. One is that the division number of the furnace and reactor is not enough in order to simplify the calculation of model. The other one is that some unknown impurities exist in the EDC feedstock in the actual process. The purity of the feed is kept at the level of 99.5% in industry, while the purity of the EDC feed is assumed at 100% in the basic simulation of the EDC cracker. 3.2. Basic Run Length Simulation Results. Due to coke deposition, the EDC cracking performance decreases with the run time when the EDC feedstock flow rate and fuel gas flow rate are kept constant. Details of the simulation results based on the simulated basic run length are presented as follows. Figure 5 and Figure 6 provide the 3D plots of coke thickness and flow diameter in the reactor as a function of the length of

Figure 5. Coke thickness along with the reactor tube length and run time.

reactor and run time, respectively. Figure 5 shows that the distribution of coke thickness is not uniform along the reactor length. The maximum coke deposition accumulates at the outlet of reactor tube. This phenomenon could be attributed to the endothermic EDC cracking reactions in the reactor. EDC feedstock is heated up mainly to increase the temperature in the first half of the reactor. No cracking reaction occurs before the temperature of the process gas reaches a certain temperature

3. RUN LENGTH SIMULATION OF AN EDC CRACKER 3.1. Model Validation. The run length of an industrial MITSUI EDC cracker is simulated in this section. Before that, the one-dimensional plug flow reactor model and Lobo−Evans furnace model should be validated first with industrial data. Table 4 summarizes some simulated crucial performance indices including the product yields at the reactor outlet vs 17507



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length gradually increases with an obvious gradient. The process gas temperature in the first half of the reactor is relatively low. With the continuous heat absorption of the process gas, its temperature rises, and this increases the external wall temperature along the reactor length. At the start of the EDC cracker operation cycle, the external wall temperature of the inlet is 664.8 K, compared with that of the outlet, which is 867.1 K. The coke thickness increases with running time, particularly at the outlet, where the external wall temperature of the reactor rises. The maximum temperature at the external wall, which is located at the reactor outlet, increases from 867.1 to 899.7 K in 70 weeks. The maximum external wall temperature limit of a reactor is the main factor in the operation cycle, as shown above. With safety considerations, the value of the maximum external wall temperature limit is conservatively set to 900 K following the recommendations of the site operator, which leads to the shutdown of the full running time of the EDC cracker for decoking after 70 weeks in this simulation. This finding is in good agreement with the actual operation cycle on site (65 weeks) and that validates the run length model of EDC cracker in certain. Full sidewall burners are adopted in the model to provide the heat in the furnace in four rows with the fuel gas allocation ratio at 18.5:18.5:27.0:36.0 from top to bottom. Figure 8 shows the

Figure 6. Flow diameter changes with the reactor tube length and run time.

(∼400 °C). The endothermic pyrolysis process starts in the last half of the reactor. With the increased concentration of the coke precursor C2H2 and the increased temperature of process gas along the reactor length, coke reaction occurs and becomes increasingly tremendous. Figure 5 depicts that coke thickness at the reactor reaches 0.0078 m when the running time reaches 70 weeks, which reduced the flow diameter at the reactor outlet from the initial 0.1013 to 0.0857 m after 70 weeks of continuous operation, as depicted in Figure 6. Coke deposition could lead to many adverse effects on the performance of the EDC pyrolysis. The main direct impact is reflected in the external wall temperature of the reactor tube. Given the poor heat transfer performance of the coke layer, thermal resistance increases with increased coke layer thickness, which forces the external wall temperature to increase with running time. The maximum external wall temperature occurs at the outlet of the reactor tube, which is consistent with the position of the maximum coke layer thickness. When the maximum external wall temperature reaches the maximum temperature limit of the metal material, the EDC cracker should be shut down for decoking, which would affect the capacity and cost of VCM production. Thus, the run time length of the cracker is directly determined by the maximum external wall temperature. The 3D plot of the external wall temperature as a function of reactor length and time is shown in Figure 7. The external wall temperature along the reactor

Figure 8. Flue gas temperature distribution throughout the furnace partitions in an operation cycle.

3D plot of flue gas temperature distribution as a function of the furnace partitioning and time. Given that the fuel gas allocation in the lower part of the furnace is larger than that of the upper part, the flue gas temperature profile along the furnace height has obvious gradient information. At the start of the running time, the flue gas temperature at the bottom is 1052.1 K, compared with that in the exit of the radiation section, which is 940.0 K. As the run time of the EDC cracker proceeds, the coke thickness increases, and results in a more difficult heat transfer. A strong heat coupling phenomenon exists between the furnace and the reactor. Thus, the flue gas temperature increases with time, particularly at the bottom of the furnace, where most coke is deposited in the corresponding position of the reactor. The flue gas temperature at the bottom rises from 1052.1 to 1058.0 K during the entire operation cycle, compared with that in the exit of the radiation section, which rises from 940.0 to 941.9 K. Figure 9 depicts the 3D plot of the heat flux profile distribution as a function of reactor length and time. With the increase in run time, the heat flux on the first several passes of the coil does not decrease. Instead, the heat flux slightly increases as a result of the low local coke rate and the increase

Figure 7. External skin temperature changes along the reactor tube length in an operation cycle. 17508



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increase in the process gas temperature in the last part of the reactor along the reactor length originally. Furthermore, as onedimensional plug flow model is assumed to describe the reactor, radial temperature difference of process gas is ignored in the reactor, which leads that the coke deposition in the bends has no big difference comparing with that in the related horizontal tubes shown in Figure 5. Figure 10 depicts that the process gas temperature increase smoothly along the reactor length and highest process gas temperature is located at the reactor outlet, which leads to major coke deposition located at the end of the longer reactor tube. Figure 10 also emphasizes that the coil outlet temperature (COT) changes from the initial 764.7 to 770.3 K at the end of the run time. The COT on site is regularly controlled at 760 to 770 K in the entire operation cycle, thereby indicating that the simulated COT is close to the industrial data. Figure 11 depicts the 3D plot of process gas pressure distribution as a function of reactor length and time. In the

Figure 9. Heat flux profile changes along the reactor tube length in an operation cycle.

in local flue gas temperature with the run time. On the contrary, the heat flux profile located in the last part of the reactor tube decreases with time because of the additional resistance to heat transfer by the coke layer, even though the flue gas temperature in the relevant position obviously increased with time. Figure 9 shows that the heat flux profile at the reactor outlet decreases from the initial 29012 W/m2 to 26029 W/m2 at the end of the operation cycle. Figure 10 illustrates the 3D plot of process gas temperature distribution as a function of reactor length and time. Process

Figure 11. Process gas pressure changes along the reactor tube length in an operation cycle.

simulation, the COT pressure was kept constant at 2400 kPa, as measured in the industrial plant. Given the formation of a coke layer on the internal tube skin, the internal cross section of the tubes decreases, which results in a higher pressure drop over the reactor length. The inlet pressure has to increase with time to maintain the required outlet pressure. The process gas inlet pressure is 3015.4 kPa at the start of the run and increased to 3109.0 kPa at the end of the run. Given the increase in pressure drop and the resulting reduction in residence time, EDC cracking performance indices decline with the run time, even though the process gas temperature slightly increased. Thus, the increase in pressure drop is the main factor in the decline of the EDC cracking performance. Figure 12 shows that the EDC cracking performance indices change with run time in an operation cycle. Figure 12a depicts the molar flow rates of the main compositions (EDC/VCM/ HCl) in the process gas outlet. The EDC molar flow rate in the process gas outlet increases from 90.4729 kmol/h at the start of the run to 92.7771 kmol/h at the end of the run, compared with that of the VCM/HCl in the process gas outlet, which decreases from 120.0062 kmol/h/120.7925 kmol/h at the start of the run to 117.7947 kmol/h/118.5478 kmol/h at the end of the run, respectively. More intuitively, the EDC cracking conversion is shown in Figure 12b. With the increase in run time, EDC conversion decreases from 0.5735 to 0.5627 in the entire operation cycle. Moreover, the EDC cracking selectivity is shown in Figure 12c. With the increase in process drop, the

Figure 10. Process gas temperature changes along the reactor tube length in an operation cycle.

gas is heated up rapidly in the first half pass of the reactor with no cracking reactions. The process gas temperature in the last half of the reactor increases slowly because most of the heat is used for EDC thermal cracking. Process gas temperature does not decrease but slightly increases with the run time even though the thermal efficiency of the furnace decreases because of the coke deposition. This phenomenon can be attributed to the increase in the overall flue gas temperature profiles. No coke deposits are observed in the first half of the reactor. Thus, the process gas temperature in the first half of the reactor increases more quickly with run time. In the last half of the reactor, the coke deposition mainly caused negative impact on the EDC cracking conversion but not on the process gas temperature. This observation is attributed to the gradual 17509



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Figure 12. (a) EDC/VCM/HCl molar flow rates in the reactor output, (b) EDC cracking conversion, (c) EDC cracking selection, (d) fuel gas consumption per unit VCM production during an operation cycle.

synergic effect on undesirable coke formation. Thus, adding the proper amount of promoter is necessary for the performances of the process based on the reaction kinetics expressions and mechanisms. The rich Cl elementary in CCl4 makes it an extremely important EDC cracking promoter. Thus, this section aims to explore the influence of CCl4 concentration on the EDC cracking process and coke formation. Simulated conditions with the concentrations of CCl4 at 0, 50, 100, 150, and 200 ppm wt % were performed. The results are given and discussed below. The promoter CCl4 is not only conducive for the conversion of EDC to VCM but can also accelerate undesirable coke formation, which is key in the operation cycle of the EDC cracker. Figure 13 shows the actual internal diameter variations of the reactor outlet with running time, which is caused by coke deposition. Figure 13 shows that lower CCl4 concentrations decreases the coke rate, which leads to a smaller gradient information in the actual internal diameter variation with running time. On the contrary, higher CCl4 concentration exacerbates the coke deposition and shortens the operation cycle of the EDC cracker, which is directly determined by the maximum reactor external temperature. Figure 14 shows the maximum reactor external skin temperature variation with runtime at different CCl4 concentrations. With the increase in CCl4 concentration, the gradient information of the maximum temperature of the external skin increases, which decreases the time for the temperature to reach the maximum temperature limit of the metal (900 K). This result implies a shorter operation time for the EDC cracker. Figure 14 depicts

residence time decreases with run time. This result reduces the EDC cracking severity but also improves the EDC cracking selectivity. Therefore, EDC cracking selectivity increases from 0.9864 to 0.9869 in the entire operation cycle. The thermal efficiency of the EDC cracking furnace declines with run time because of coke deposition. With the constant supply of fuel gas and the continuous decline of EDC cracking performance, the flue gas consumption per unit VCM production, which represents the cost of VCM production in the industry, increases with run time, as shown in Figure 12d. Figure 12d shows that the fuel gas consumption per unit VCM production increases from 66.6166 kg/t at the start of the run to 67.8673 kg/t at the end of the run. From Figure 5, it can be clearly seen that coke layer thickness increases approximately linearly with time increase, especially more obvious at the reactor outlet, resulting in linear increase of external skin temperature and linear reduction of heat profiles at the reactor outlet. The factors mentioned above lead to that the key performance indicators depicted in Figure 12 change approximately linearly.

4. DISCUSSION ON THE ADDITION OF PROMOTER CCL4 One of the factors that affect EDC cracking performance is the composition of the raw material, which is very critical. Great care should be taken to ensure that the EDC used for cracking to VCM is of high purity. However, minimal trace amounts play an important role in EDC cracking inhibition and fouling, which can act either as promoter or inhibitor and have a 17510



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time at various CCl4 concentrations. Figure 15 shows that the EDC conversion at the beginning of the run-time is 0.5756/ 0.5787/0.5838/0.5859/0.5910 at concentrations of 0, 50, 100, 150, and 200 ppm wt %, respectively. As the running time proceeds with the increase in the thickness of the coke layer, heat-transfer efficiency decreases, the pressure drop of the process gas increases, and the residence time of the process gas is shortened. All the factors mentioned above decreases the conversion of the EDC cracking with time. By the end of the operation cycle in each case, the conversion decreased to 0.5628/0.56813/0.57335/0.5786/0.5841. Figure 16 shows the EDC cracking selectivity variations with run time at different CCl4 concentrations. The addition of

Figure 13. Actual internal diameter variations of reactor output location with different CCl4 concentrations.

Figure 16. EDC cracking selection variations with different CCl4 concentrations during the operation cycle.

CCl4, on the one hand, can increase the conversion of EDC cracking. On the other hand, it also increases the production of byproducts in the EDC cracking. Increasing the concentration of CCl4 reduces the EDC cracking selectivity. Figure 16 shows that the EDC selectivity for the clean tubes with the five different CCl4 concentrations are 0.9864, 0.9831, 0.9799, 0.9767, and 0.9736, respectively. The conversion decreases with the increase in running time, whereas the selectivity increases. At the end of each operation cycle at the five conditions, the selectivity increases to 0.9869, 0.9836, 0.9803, 0.9771, and 0.9739, respectively. The promoter CCl4 induces the conversion of EDC to VCM and HCl under the conditions of a fixed amount of fuel gas supply, which implies that increasing the concentration of CCl4 improves the VCM production. Thus, the fuel gas consumption per unit of VCM production is reduced. Figure 17 shows the fuel gas consumption per unit of VCM production variations with running time with five different amounts of CCl4. At the initial running stage of the EDC cracker, the fuel gas consumption per unit of VCM production with the five levels of CCl4 concentration are 66.61, 66.24, 65.87, 65.51, and 65.15 kg/t, respectively. As the running time proceeds and as a result of the coke deposition, the heat efficiency of the cracker is reduced, and the residence time of the process gas becomes shorter. Thus, the cost per unit of VCM production increases during the operation. At the end of the EDC cracker operation cycle of each case, the fuel gas consumption per unit of VCM production reaches 67.85, 67.44, 67.05, 66.66, and 66.24 kg/t, respectively.

Figure 14. Maximum reactor external skin temperature variations with different CCl4 concentrations.

operating cycles of 70, 65, 61, 57, and 52 weeks at CCl4 concentrations of 0, 50, 100, 150, and 200 ppm % wt, respectively. The promoter CCl4 is rich in elementary Cl. Thus, EDC cracking is greatly improved with the addition of CCl4, which is conducive for the conversion of EDC to VCM and HCl. Figure 15 shows the EDC cracking conversion variations with running

Figure 15. EDC cracking conversion variations with different CCl4 concentrations during the operation cycle. 17511



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Figure 17. Fuel gas consumption (per unit VCM production) variations with different CCl4 concentrations.

Figure 18. Maximum reactor external skin temperature variations with different fuel gas allocation strategies.

In conclusion, high-purity raw material is required in the EDC cracking process. Particularly, the concentration of the cracking promoters should be controlled precisely in the ppm level. Although excessive addition of promoter can improve EDC conversion, this setup can also accelerate the coke rate in the EDC cracking process, thereby shortening the operating cycled. The optimization of CCl4 concentration in the raw material is crucial in improving the overall economic benefits of the cracker. On the bais of economic benefit analysis and onsite implementation experience, the CCl4 concentration should be maintained at 100 ppm wt % to ensure high conversion rates and guarantee that the run cycle is maintained at a reasonable level such that the cracker gains the most benefits during the entire operation cycle.

the consistency of the maximum reactor external skin temperature and process gas temperature. The maximum external temperature occurs at the reactor outlet. Thus, increasing the fuel gas allocation factor inevitably leads to an increase in flue gas temperature at the bottom. Therefore, the external skin temperature at the reactor outlet (maximum temperature) is also increased. As the running time proceeds, the maximum external skin temperature of the reactor increases because of coke deposition. The gradient information is not similar among the six fuel gas allocation strategies. Given that the coke reaction mainly occurs at the end of the reactor tube, an increase in fuel gas allocation factor causes an increase in coke rate, which results in the rapid increase in external skin temperature at the end of the reactor. This setup leads to a shorter operation cycle of the EDC cracker. When the maximum reactor external skin temperature reaches the limit temperature due to the metallurgy, the cracking furnace should be shut down for decoking. Figure 18 shows that the run lengths with the six fuel gas allocations are 139, 110, 88, 70, 56, and 44 weeks, respectively. Figure 19 depicts the actual flow diameter variations at the reactor outlet with running time under the different fuel gas allocations. As the running time proceeds, the thickness of the coke layer gradually increases such that the inner diameter of the reactor tube outlet is gradually reduced. A greater fuel gas allocation factor causes a larger decrease in gradient

5. DISCUSSION OF FUEL GAS ALLOCATION Apart from the purity of raw material, fuel gas allocation in the furnace also has a crucial influence on the EDC cracking performance under the condition of a fixed fuel gas amount. In this paper, the fuel gas allocation factor α is defined as the ratio of the fuel gas flow in the last burners’ line to the total fuel gas flow in the furnace according to the fuel gas control program applied on site, as depicted in Figure 4. Different fuel gas allocation strategies lead to different combustion conditions, which cause different flue gas temperature distributions within the furnace. Given the severe thermal coupling between the furnace and the reactor tube, if more fuel gas is allocated in the bottom of the furnace, much more heat would be released and transferred to the end of the According to the actual on-site operation, the fuel gas allocation factor can only be adjusted within a small range (0.3 to 0.4), in which six simulation points are taken at 0.30, 0.32, 0.34, 0.36, 0.38, and 0.40. Figure 18 shows the maximum reactor external skin temperature variation with running time using various fuel gas allocation strategies. Different fuel gas allocation strategies cause different flue gas temperature distributions in the furnace, which leads to various external skin temperature distributions. Given the intense thermal coupling between the furnace and the reactor tube, as shown in Figure 18, the maximum reactor external skin temperature with the six fuel gas allocation factors at the start of the operation cycle are 855.81, 859.67, 863.45, 867.12, 870.73, and 874.23 K, respectively. With the increase in the fuel gas allocation factor α, most of the combustion reactions are concentrated at the furnace bottom according to

Figure 19. Actual internal diameter variations of reactor output location with different fuel gas allocations. 17512



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information of flow diameter. When the end of the operating cycle of each condition is reached, the actual flow diameter at the reactor outlet is 0.07886, 0.0812, 0.0835, 0.0857, 0.08761, and 0.08951 m, respectively. Figure 20 presents the reactor coil temperature variation with running time under different fuel gas allocations. The increased

Figure 21. EDC cracking conversion variations with different fuel gas allocation strategies.

reduces the EDC cracking selectivity at the same time. Figure 22 shows the EDC cracking selectivity variations with running

Figure 20. COT variations with different fuel gas allocation strategies.

fuel gas allocation at the bottom induces the heat transfer between the furnace and reactor. Thus, the process gas temperature at the reactor outlet is improved by increasing the fuel gas allocation factor. Figure 20 shows that at the beginning of the run, COT values are 762.61, 763.68, 764.67, 765.60, and 766.48 K for the six various fuel gas allocations, respectively. With the increase in running time, the thickness of the coke layer gradually increases, the pressure drop in the pipe increases, and the residence time becomes shorter, which reduces the heat efficiency of the cracker and the EDC cracking performances. However, a greater proportion of the heat absorbed by the process gas is used for the increase in process gas temperature. Thus, COT increases over time. A higher fuel gas allocation factor corresponds to more severe coke deposition, which leads to a higher increase in COT. At the end of the running cycle for each condition, COT is 768.28, 768.81, 769.32, 769.72, 770.12, and 770.43 K, respectively. When the fuel gas allocation factor α increases, most of the fuel gas is combusted at the furnace bottom. Thus, the heat exchange of the flue gas within the furnace is maximized and the thermal efficiency of the furnace is improved, which shows that more heat is provided for the reactor, particularly at the reactor end. The EDC cracking conversion is improved by the increase in fuel gas allocation factor. Figure 21 shows the EDC cracking conversion variations with running time for the six various fuel gas allocations. These results indicate that under clear tube condition, the EDC cracking conversions are 0.5703, 0.5712, 0.5723, 0.5736, 0.5750, and 0.5766, which correspond to the six fuel gas allocation factors of 0.30, 0.32, 0.34, 0.36, 0.38, and 0.40, respectively. With the increase in run time, the cracking performance decreases with coke deposition. A high fuel gas allocation factor leads to more serious coke deposition, which reduces the EDC conversion more quickly. Figure 21 shows that the conversions at the end of each operation cycle are 0.5557, 0.5579, 0.5603, 0.5628, 0.5654, and 0.5681, respectively. Improving the fuel gas allocation factor, on one hand, is conducive for improving EDC conversion. On the other hand, it also increases the side reaction rates of EDC cracking and

Figure 22. EDC cracking selectivity variations with different fuel gas allocation strategies.

time under the conditions of the different fuel gas allocation strategies. At the start of running the EDC cracker, the selectivity under different fuel gas allocations are 0.986459, 0.986434, 0.986402, 0.986361, 0.986313, and 0.986259, respectively. With the increase in running time, the increase in thickness of the coke layer reduces the heat efficiency of the EDC cracker; and the EDC cracking conversion decreases with time. The selectivity, which exists on the opposite side of the conversion, increases with running time. Figure 22 shows that a higher fuel gas allocation factor corresponds to a more obvious increased rate of selectivity. At the end of each operation cycle in the six cases, the EDC cracking selectivity reaches 0.987126, 0.987052, 0.986965, 0.986870, 0.986772, and 0.986664, respectively. In the case of a fixed total amount of fuel gas, increasing the fuel gas allocation factor within the reasonable range improves the conversion from EDC to VCM, which implies that the fuel gas consumption per unit of VCM production is reduced by increasing the fuel gas allocation factor. Figure 23 shows the fuel gas consumption variations per unit of VCM production throughout the running time under the six conditions of different fuel gas allocation strategies. At the start of running the EDC cracker, the values of the fuel gas consumption per unit of VCM production are 66.99, 66.88, 66.75, 66.61, 66.45, and 66.27 kg/t, which correspond to the six fuel gas allocation 17513



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fuel gas allocation factor should be maintained at 0.36 to guarantee the overall economic benefits of the EDC cracker in the full operation cycle. Generally speaking, it is a typical multiconstrained multiobjective optimization problem making a good trade-off among the core cracking performance indices. But the calculation of the run length model of thermal coupled EDC reactor/furnace is time-consuming, which makes evolution algorithms, such as the differential evolution algorithm or genetic algorithm, unsuitable for EDC cracker optimization at present stage. However, this work lays a solid foundation for the further multiobjective optimization research based on the core performance indices of the EDC cracker.

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Figure 23. Fuel gas consumption (per unit VCM production) variations with different fuel gas allocation strategies.

AUTHOR INFORMATION

Corresponding Authors

*E-mail: [email protected]. *E-mail: [email protected].

factors. As the running time proceeds, the EDC cracking performances generally decrease and the fuel gas consumption per unit of VCM production increases because of coke deposition. A higher fuel gas allocation factor corresponds to a more serious increase in fuel gas consumption. At the end of the each operating cycle in the six cases, the fuel gas consumption per unit VCM production increases to 68.69, 68.43, 68.15, 67.85, 67.55, and 67.23 kg/t, respectively. In conclusion, the fuel gas allocation has a very critical influence on the heat transfer of EDC cracker, EDC cracking performance indices, and operation cycles. Increasing the proportion of fuel gas allocation at the bottom of the furnace is beneficial to the improvement of the EDC cracking conversion, but it reduces the selectivity while causing severe coke deposition, which greatly shortens the run cycle. Therefore, choosing a suitable fuel gas ratio is very critical to the overall economic benefits in the entire operation cycle of the EDC cracker. The fuel gas allocation factor α should be maintained at 0.36 to ensure a long run cycle of the EDC cracker and a high conversion average.

Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS This work is supported by Major State Basic Research Development Program of China (2012CB720500), National Natural Science Foundation of China (U1162202, 61174118, 61222303, 21276078), and Shanghai Leading Academic Discipline Project (B504).

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6. CONCLUSION The full cycle of an industrial EDC cracker is simulated. Given the intense heat coupling between the furnace and the reactor, the simulation of the EDC cracker is divided into two parts: the furnace model simplified with one-dimensional Lobo−Evans method and the reactor with one-dimensional plug flow model. Radical mechanism is adopted to describe the EDC cracking reactions with 24 reaction equations and 31 components. On the basis of the full cycle simulation, two important factors, namely, CCl4 concentration and fuel gas allocation, are investigated to improve the overall benefits of the whole operation cycle, including conversion, selectivity, fuel gas consumption per unit of VCM production, and run length. The addition of EDC cracking promoter CCl4 can improve the EDC conversion, but it can also aggravate the coking reaction, which causes the sharp deterioration of the cracker performance and shortening of the running cycle. Increasing the fuel gas proportion at the furnace bottom can effectively improve the heat transfer efficiency of the EDC cracker. In particular, this process enhances the heat transfer at the end of the reactor tube. Thus, EDC conversion can be improved, but due to coke deposition, the run cycle is obviously shortened. Through sensitive analysis, the concentration of the promoter CCl4 should be controlled at the level of 100 ppm wt % and the 17514

NOMENCLATURE aD = angle factor Acpi = cold plane, m2 ARi = tubes’ skin surface in the ith zone, m2 A0 = effective axial radiation surface, m2 cpj = heat capacity of the jth species, J/mol/K Cp = heat capacity, J/kg/K Cpi = heat capacity of pure species i, J/kg/K dt,dt_new,dt_old = the diameter of the reactor tube, m di = inner diameter, m do = outer diameter, m Fj = molar flow rate, mol/s Fr = friction factor Fi = radiation heat transfer coefficient G = total mass flow rate, kg/m2 /s ΔH = heat of reaction, J/mol ΔHi = net heat of flue gas enthalpy change, W K = total heat transfer coefficient, W/(m2.K) Mm = coke molecular weight, g/mol Mm = average molecular weight of all the process gas species, g/mol Mi,Mj = molecular weight of species i,j, g/mol Pt = total pressure, Pa Q = heat flux, W/m2 Qfl,Qfli = low heating value, J/kg Qt = heat transfer, W Qsi = heat release of fuel gas in the local zone, W Qri = heat absorbed by the reactor tubes located in ith zone, W Qli = the furnace wall heat dissipation in the ith zone W R = universal gas constant, J/mol/K rri = reaction rate, mol/m3/s Re = Reynolds number Rb = radius of the tube bend, m dx.doi.org/10.1021/ie401265f | Ind. Eng. Chem. Res. 2013, 52, 17501−17516


Article

rc(z) = Coking rate, kg/m2 /s Ri = inner fouling resistance, m2.K/W Ro = reactor outer tube wall fouling resistance, m2.K/W sij = stochiometry factor T = local temperature and process gas temperature, K Two = reactor outer wall temperature, K Tgmi = average temperature of combustion in the ith zone, K Twi = average temperature of tube outer skin in the ith zone, K Xi = mole fraction of species i YiYfi = mass fraction of species i z = location variable, m

(11) Mochida, I.; Tsunawaki, T.; Sotowa, C.; Higuchi, K. Coke Produced in the Commercial Pyrolysis of Ethylene Dichloride into Vinyl Chloride. Ind. Eng. Chem. Res. 1996, 35 (10), 3803−3807. (12) Borsa, A. G.; Herring, A. M.; McKinnon, J. T.; McCormick, R. L.; Yamamoto, S.; Teraoka, Y.; Natori, Y. Characterization of Coke Formed in Vinyl Chloride Manufacture. Ind. Eng. Chem. Res. 1999, 38 (11), 4259−4267. (13) Borsa, A. G.; Herring, A. M.; McKinnon, J. T.; McCormick, R. L.; Ko, G. H. Coke and Byproduct Formation during 1,2-Dichloroethane Pyrolysis in a Laboratory Tubular Reactor. Ind. Eng. Chem. Res. 2001, 40 (11), 2428−2436. (14) Plehiers, P. M.; Reyniers, G. C.; Froment, G. F. Simulation of the Run Length of an Ethane Cracking Furnace. Ind. Eng. Chem. Res. 1990, 29 (4), 636−641. (15) Heynderickx, G. J.; Froment, G. F.. Simulation and Comparison of the Run Length of an Ethane Cracking Furnace with Reactor Tubes of Circular and Elliptical Cross Sections. Ind. Eng. Chem. Res. 1998, 37 (3), 914−922. (16) Yancheshmeh, M. S. S.; Haghighi, S. S; Gholipour, M. R.; Dehghani, O.; Rahimpour, M. R.; Raeissi, S.. Modeling of Ethane Pyrolysis Process: A Study on Effects of Steam and Carbon Dioxide on Ethylene and Hydrogen Productions. Chemical Engineering Journal 2013, 215, 550−560. (17) Berreni, M.; Wang, M. Modelling and Dynamic Optimization of Thermal Cracking of Propane for Ethylene Manufacturing. Comput. Chem. Eng. 2011, 35 (12), 2876−2885. (18) Gao, G. Y.; Wang, M.; Ramshaw, C.; Li, X. G.; Yeung, H.. Optimal Operation of Tubular Reactors for Naphtha Cracking by Numerical Simulation. Asia-Pac. J. Chem. Eng. 2009, 4 (6), 885−892. (19) Zychlinski, W.; Boettger, G.; Rodewald, D. Coke Formation in Thermal Decomposition of 1,2-Dichloroethane: Conclusions from Radiotracer Experiments. Chem. Ing. Tech. 1990, 42 (8), 325 (in German). (20) Albright, L. F.; Marek, J. C.. Coke Formation During Pyrolysis: Roles of Residence Time, Reactor Geometry, and Time of Operation. Ind. Eng. Chem. Res. 1988, 27 (5), 743−751. (21) Graff, M. J.; Alberight, L. F.. Coke Deposition from Acetylene, Butadiene and Benzene Decompositions at 500−900 °C on solid surfaces. Carbon 1982, 20 (4), 319−330. (22) Kopinke, F. D.; Zimmermann, G.; Nowak, S.. On the Mechanism of Coke Formation in Steam CrackingConclusions from Results Obtained by Tracer Experiments. Carbon 1988, 26 (2), 117−124. (23) Kopinke, F. D.; Zimmermann, G.; Reyniers, G.; Froment, G. F. Relative Rates of Coke Formation from Hydrocarbons in Steam Cracking of Naphtha. 2. Paraffins, Naphthenes, Mono-, Di-, and Cycloolefins, and Acetylenes. Ind. Eng. Chem. Res. 1993, 32 (11), 56− 60. (24) Kopinke, F.-D.; Zimmermann, G.; Reyniers, G.; Froment, G. F. Relative Rates of Coke Formation from Hydrocarbons in Steam Cracking of Naphtha. 3. Aromatic Hydrocarbons. Ind. Eng. Chem. Res. 1993, 32 (11), 2620−2625. (25) Palmer, H. B.; Cullis, C. F. Chemistry and Physics of Carbon; Marcel Dekker Inc.: New York, 1965. (26) Ranzi, E.; Dente, M.; Pierucci, S.; Barendregt, S.; Cronin, P. Coking Simulation Aids On-Stream Time. Oil Gas J. 1985, September (2), 49−52. (27) Lacroix, R.; Fournet, R.; Ziegler-Devin, I.; Marquaire, P.M.. Major Precursors of Pyrocarbon Deposition from Propane Pyrolysis. ECS Trans. 2009, 25 (8), 91−98. (28) Albright, L. F.; Marekt, J. C.. Mechanistic Model for Formation of Coke in Pyrolysis Units Producing Ethylene. Ind. Eng. Chem. Res. 1988, 27 (5), 755−759. (29) Niaei, A.; Towfighi, J.; Sadrameli, M.. Prediction of Furnace Run Length for the Pyrolysis of Naphtha by a PC Based Simulator. Sci. Iran. 2003, 10 (3), 287−299. (30) Plehiers, P. M.; Reyniers, G. C.; Froment, G. F. Simulation of the Run Length of an Ethane Cracking Furnace. Ind. Eng. Chem. Res. 1990, 29, 636−641.

Greek Letters

α = fuel gas allocation factor, coke laydown factor and convectional heat transfer coefficients η = conversion factor, atm Pa−1 ζ = parameter of tube bend Λ = angel of bend δ(z) = local coke thickness, m Δδc(z) = local coke thickness increment, m ρc = density of the coke, kg/m3 λ = thermal conductivity, W/m/K λc = thermal conductivity of coke layer, W/m/K λw = thermal conductivity of reactor tube wall, W/m/K λi = thermal conductivity of pure species i, W/m/K σ = Stefan−Boltzmann constant, σ = 5.672 × 10−8 W/m2 K4 σi = Lennard-Jones collision diameter of species i Å ϕij = interaction parameter μi μj = viscosity of gas molecules, kg/(m/s) Ωμi = viscosity collision integration times of species i

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REFERENCES

(1) Li, C.; Hu, G.; Zhong, W.; Cheng, H.; Du, W.; Qian, F.. Comprehensive Simulation and Optimization of an Ethylene Dichloride Cracker Based on the One-Dimensional Lobo−Evans Method and Computational Fluid Dynamics. Ind. Eng. Chem. Res. 2013, 52, 645−657. (2) Barton, D. H. R.. Kinetics of the Dehydrochlorination of Substituted Hydrocarbons. Part I. Induced Dehydrochlorination. J. Chem. Soc. 1949, 148−155. (3) Barton, D. H. R.; Howlett, K. E.. Kinetics of the Dehydrochlorination of Substituted Hydrocarbons. Part II. The Mechanism of the Thermal Decomposition of 1:2-Dichloroethane. J. Chem. Soc. 1949, 156−164. (4) Huybrechts, G.; Wouters, G.. Mechanism of the Pyrolysis of 1,2Dichloroethane in the Absence and Presence of Added Chlorine. Int. J. Chem. Kinet. 2002, 34 (5), 316−321. (5) Ashmore, G. P.; Owen, A. J.; Robinson, P. J.. Chlorine-Catalysed Pyrolysis of 1,2-Dichloroethane. Part 1. Experimental Results and Proposed Mechanism. J. Chem. Soc., Faraday Trans. 1982, 78, 657− 676. (6) Ranzi, E.; Dente, M.; Rovaglio, M.. Pyrolysis and Chlorination of Small Hydrocarbons. Chem. Eng. Commun. 1992, 17 (1), 17−39. (7) Borsa, A. G.. Industrial Plant/Laboratory Investigation and Computer Modeling of 1,2-Dichloroethane Pyrolysis. Ph.D. Dissertation, Colorado School of Mines, Golden, CO, 1999. (8) Song, B. H.; Choi, B. S.; Kim, S. Y.; Yi, J.. Kinetics and Mechanism of Pyrolysis of the 1,2-Dichloroethane. J. Korean Inst. Chem. Eng. 2001, 99, 481−487. (9) Choi, B. S.; Oh, J. S.; Lee, S. W.; Kim, H.; Yi, J.. Simulation of the Effects of CCl4 on the Ethylene Dichloride Pyrolysis Process. Ind. Eng. Chem. Res. 2001, 40, 4040−4049. (10) Schirmeister, R.; Kahsnitz, J.; Trager, M.. Infuelnce of EDC Cracking Severity on the Marginal Costs of Vinyl Chloride Production. Ind. Eng. Chem. Res. 2009, 48, 2801−2809. 17515



Article

(31) Sundaram, K. M.; Froment, G. F. Kinetics of Coke Deposition in the Thermal Cracking of Propane. Chem. Eng. Sci. 1979, 34 (5), 635−644.

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