Thermal Cracking of Hydrocarbon Aviation Fuels in Regenerative

Article pubs.acs.org/EF

Thermal Cracking of Hydrocarbon Aviation Fuels in Regenerative Cooling Microchannels Rongpei Jiang,† Guozhu Liu,*,† and Xiangwen Zhang Key Laboratory for Green Chemical Technology of Ministry of Education, School of Chemical Engineering and Technology, Tianjin University, Tianjin 300072, People’s Republic of China S Supporting Information *

ABSTRACT: Regenerative cooling with hydrocarbon aviation fuels on board is taken as a promising technology for the thermal management system of next-generation aircraft. An improved methodology of an electrically heated tube (1 mm i.d.), i.e., applying the variable reactor tube length to carry on thermal cracking of supercritical hydrocarbon aviation fuels as the electric current heating maintains constant, was proposed to experimentally obtain detailed information on the local concentration and temperature along the microchannels of a heat exchanger. For the first time a series of experimental data on detailed local chemical compositions of cracked hydrocarbon fuel along the cooling microchannels were reported under supercritical conditions (5 MPa, 680−700 °C), and the calculated thermodynamic properties, velocity, and residence times along the tube were also reported. A modified molecular reaction model consisting of 18 species and 24 reactions was developed to predict thermal cracking of hydrocarbon aviation fuels in a wide range of cracking conversion (up to 86%). The work is significant for the design of regenerative cooling structures in predicting the local chemical compositions, estimating thermophysical properties, and coking of the cracked hydrocarbon fuels for heat transfer analysis.

1. INTRODUCTION Nowadays, hypersonic aircraft have been attracting extensive attention due to increasing interesting in high-speed flight.1−4 When the flight speed of future aircraft reaches or exceeds Mach 5 (ca. 1700 m/s), one of the most important challenges is the exposure of some aircraft components (e.g., combustor) to such high aerothermal loads that classical materials such as metals or simple ceramics can no longer be used.5 Regenerative cooling by using hydrocarbon fuel on board has been considered to be the most effective solution,6 wherein liquid hydrocarbon fuel flows through parallel microchannels (equivalent diameter of 1−2 mm) of heat-exchanger panels outside the combustor before being injected into the combustor (see Figure 1). In this case, hydrocarbon fuels may be heated even up to 750 °C to provide a heat sink capacity of more than 3.5 MJ/kg due to “sensible” heating and endothermic chemical reactions (e.g., thermal cracking), and thus the term “endothermic hydrocarbon fuel” (EHF) is usually used for hydrocarbon fuels with bulk endothermic reactions.5−11 Furthermore, the thermal cracking of the fuels into smaller molecules (hydrogen, ethylene, etc.) is also favorable for improving the combustion of hydrocarbon fuel in the supersonic air flow by shortening ignited delay times.12 It is worth noting that hydrocarbon fuels are also exposed to high pressures of 3−8 MPa, higher than the critical pressure of kerosene fuels of ∼2 MPa, to prevent boiling of heated fuels and maintain flow in high-pressure combustors. Typically, EHFs are supercritical fluids before significant decomposition.13 The most important issue is to develop the EHFs with enough heat sink under supercritical conditions. JP-7 (MIL-DTL38219), a typical EHF, meets the operational demands for supersonic and even hypersonic aircraft (Mach 5+). Unfortunately, it is no longer produced. Nevertheless, several © 2013 American Chemical Society

candidates are being considered as the replacement of JP-7, involving various types of rocket grade kerosene (for example, RP-1 and RP-2, MIL-DTL-25576D) as well as JP-8 blended with additives.14 Furthermore, another crucial issue is to understand the coupled and complicated phenomena of the pyrolysis, fluid flow, and heat transfer of cracked EHFs for the rational design and analysis of regenerative cooling panels. To get more information about hydrocarbon fuels in the regenerative cooling channel, researchers developed several important experimental apparatus and methods, among which the electrically heated tube test (EHTT) was widely used due to its simplicity. Spadaccini and Sobel used EHTT to evaluate the heat sinks of several typical jet fuels (JP-8, JP-8+100, JP-7, JP-10).2,7 Linne et al. investigated the heat transfer and coking of JP-7 using EHTT experiments.15 Zhong et al.16 experimentally studied the heat transfer behaviors of Chinese RP-3 fuel using EHTT, and developed supercritical heat transfer correlations using the calculated thermophysical properties with the surrogate compounds. However, to the best of our knowledge, there were still no available experimental data on the local chemical compositions of cracked EHFs along the electrically heated tube because of the considerable difficulty brought by the small size (ca. 1−2 mm), regardless of its crucial importance in estimating local thermophysical properties of cracked EHFs. In addition, another challenge of the regenerative cooling, coke deposited on microchannel, is also strongly dependent on the local chemical compositions of cracked EHFs, especially the content of some coke precursors (e.g., olefins and aromatics).17,18 Consequently, a new strategy Received: March 4, 2013 Revised: March 31, 2013 Published: April 9, 2013 2563

dx.doi.org/10.1021/ef400367n | Energy Fuels 2013, 27, 2563−2577

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Figure 1. Schematic representation of regenerative cooling technology for advanced aircraft.

their model is too general to provide any information on the products.26 On the contrary, some researchers developed detailed kinetics models containing more than several thousand elemental reactions.27,28 It is obviously merited to fulfill the pyrolysis mechanism, yet the massive numbers of reactions and species will inevitably raise a great challenge. Although extensive literature has been published on the experimental and modeling aspects, the high-pressure thermal cracking kinetic models of hydrocarbon fuels (especially for blended hydrocarbon fuels, such as JP-7 and RP-3) with reliable product distributions are still not available for potential engineering applications. The objective of this work is to provide new insights into the thermal cracking behaviors of EHFs in regenerative cooling passages using a new electrically heat tube method, i.e., series test run of variable tube length with the same heating electric current. Thermal cracking of hydrocarbon aviation fuel was carried out under given conditions (0.4−1.0 g/s, 1 mm i.d., 5 MPa) to provide detailed local chemical compositions of cracked hydrocarbon fuel downstream along the cooling microchannels. After that, the local densities and real velocities of cracked hydrocarbon fuel were calculated with SUPERTRAPP software. On the basis of the calculated residence time, a molecular kinetics model was developed to describe the thermal cracking of supercritical hydrocarbon fuel. This work provided some interesting and valuable experimental results and kinetics approaches on the thermal cracking of supercritical EHFs in the regenerative cooling microchanels, which is one of the crucial issues in the rational design of regenerative cooling structures.

(or method) of EHTT is still necessary to provide more details on the thermal cracking of an EHF along the cooling channel, and thus to elucidate the effect of cracking and coking reactions on the heat transfer in the regenerative cooling channels. For developing a design tool for regenerative cooling technology, the thermal cracking kinetics of hydrocarbon fuels is another important concern to predict the chemical compositions and thus to estimate the thermophysical properties of cracked EHFs. Previously, great attempts were made toward developing a robust kinetics of thermal cracking of pure or blended hydrocarbon fuels under high pressure. Sheu et al.19,20 proposed a three-step lumped model to approximately describe the thermal cracking of Norpar-13 (a mixture of nalkanes, primarily C12−C14, without current availability because of environmental concerns) in an electrically heated tube reactor (2.16 mm in i.d.) under 6.89 MPa, in which the compositions of gas and liquid phase products were approximately taken as constant. However, their model neglected the secondary reaction between the gaseous products, posing a limit to its application at high cracking degrees. Ward et al.21,22 studied the thermal cracking of ndecane and n-dodecane in a flowing tube reactor (2.16 mm in i.d.) under 3.45−11.38 MPa. They noted that the selectivities of both light products (methane, ethane, propene, propane, 1butene, n-butane, etc.) and heavier 1-alkene products (1pentene, 1-heptene, 1-octene, 1-decene, 1-nonene, etc.) were approximately constant when the conversions were less than 20%, and then proposed a concept of proportional product distribution (PPD) and developed a one-step model for the mild cracked normal alkanes. The PPD−Ward model was effective in modeling the cracking of pure hydrocarbons and kerosene fuels, as shown by recent work of Bao et al.23 and Hou et al.24 However, the use of the PPD−Ward model was also strictly limited for conversions less than 20%, which also to some extent limited its application. Zhong and co-workers25 studied the thermal cracking of China No. 3 jet fuel (also known as RP-3), a kerosene type hydrocarbon fuel, in an electrically heated tube (12 mm in i.d.) under 3.5 MPa. They also developed a one-step lumped model with the cracked fuel grouped into three categories: unreacted kerosene, gaseous products, and liquid residuals. Zhong’s approach yet fails to incorporate the formation of aromatics (benzene, toluene, xylene, etc.) as well as other compounds in their model. Very recently, Widegren and Bruno carried out the thermal cracking of RP-1 and RP-2 in stainless-steel ampule reactors. They derived global pseudo-first-order rate constants that approximate the overall rate of decomposition for the fuel, whereas

2. EXPERIMENTAL SECTION 2.1. Materials. Two hydrocarbon fuels (HFs) were used in this work. As a typical hydrocarbon jet fuel, HF-I, a commercial Chinese No. 3 jet fuel (meeting the specification of Jet-A) after special hydrotreating, was used in the developed experimental method and the corresponding kinetics approaches. HF-II, another fuel with different hydrocarbon compositions, was used to verify the developed methods and kinetics. Several major specification properties and composition information of HF-I and HF-II are listed in Table 1. 2.2. Electrically Heated Tube Test of Hydrocarbon Fuel. An electrically heated tube was utilized (similarly to that of Huang et al.7) to simulate the heat-exchanger microchannel in which hydrocarbon fuels flow and thermally crack. Figure 2a illuminates a schematic diagram of experimental apparatus for the electrically heated tube used in this work. The experimental apparatus consisted of five subsystems, i.e., feeding system, direct current (dc) power system, electrically heated tube reactor, data acquisition/controlling system, and sampling and online analysis system. Before the experimental runs nitrogen was 2564


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The experimental uncertainty is dependent on the measurements of the electrical current used to heat, fuel entrance, and exit temperaure, and the fuel flow rate. The error of the electrical current of dc power measurement was less than 0.5%. The errors of the fuel temperature measurement were dependent on the K-type thermocouples (less than ±1.5 °C between 450 and 750 °C). The errors of the fuel flow rate were less than 0.5%. On the basis of those measurement errors, the overall experimental uncertainty of this method is approximately 3.5%. We must say that the formation of amorphous (e.g., bulk-fuel phase) carbon is fully expected for the fuel and reaction conditions (e.g., levels of conversion) used in this study. The level of deposition will be strongly affected by the time on stream during the testing. In our test, average and maximal coke amounts along the tube were less than 30 and 50 mg/cm2 after each run. In this sense, the effect of carbon deposit (and subsequently inhibition of heat transfer) during the experimental testing was negligible. 2.3. Proposed Experimental Method. The fuel in the tube is heated through resistive heating. We noted that the electrical current at any position of the tube was uniformly equivalent, and that both the local electrical resistance and heat loss were functions of the wall temperature. Therefore, for the total tube length L the heat sinks (or the endothermic heat) of the fuel could be calculated by the total heat input (I2R) minus the heat lost with the following equation:

Table 1. Specification Properties of Hydrocarbon Aviation Fuel Used in This Work property density (20 °C), g/mL freezing point, °C flash point, °C critical press., MPa critical temp, °C ASTM distillation, °C IBP 10% 20% 50% 90% FBP hydrocarbon type, wt % n-alkanes isoalkanes naphthenes aromatics total heat value, MJ/kg total sulfur, ppm formula

HF-I

HF-II

0.8030 −49 52 2.420 375.45

0.8340 −75 77 2.351 421.41

169.3 188.6 195.3 206.8 221.6 235.7

195.0 206.7 209.7 212.9 223.3 239.4

44.85 27.11 25.32 2.71 43.3 100) to guarantee a negligible terminal effect. We ran a series of experiment by heating different lengths of reactor tube to measure the wall temperature downstream distributions (80 A, 0.526 g/s, 5 MPa) and the results are plotted in Figure S-1 in the Supporting Information. For instance, all the wall temperatures at 0.3 m are 687 ± 5 °C for the heating length from 0.30 to 0.60 m, and they fall first at the initial heated section (0 < l < 0.12 m) and then increase as the tube distance extends (l > 0.12 m). A similar result was also observed by Li et al.29 The temperature downstream distributions from series runs show approximate coincidence with each other regardless of the

used to blow and replace the air in the tube. The hydrocarbon fuel (25 °C) was pumped at a given flow rate with a P230 high-pressure liquid chromatography pump (Dalian Elite Analytical instruments Co., Ltd.). An online electronic balance was also used to guarantee the accuracy of hydrocarbon fuel flow rate. The electrically heated tube reactor (2.0 mm o.d. and 1.0 mm i.d.) was made of a nickel-based superalloy (for more details, see Table S-1 in the Supporting Information) coated with an Al2O3 coating of 0.5 μm to inhibit the surface disposition by deactivating the surface reactivity.18 A dc stabilized power supply was used to provide direct current voltage with two copper bars outside the tube. The hydrocarbon fuel was resistively heated and underwent thermal cracking reactions in the tube. The cracked fuel first flowed into a water-cooled heat exchanger to be quenched to 30 °C, and then into a gas−liquid separator after a Tescom back-pressure valve (Emerson Electric Co., USA) which was used to maintain the system pressure at 5.0 MPa. The inlet and outlet temperatures of fuel were measured by K-type thermocouples inserted into union cross junctions. The wall temperatures were measured by K-type thermocouples welded outside the tube at 5 cm intervals. The system pressure was detected with two Rosemount pressure transducers. This resistive heating method allowed direct measurement of the overall heat sink capacity of the fuels by performing an energy balance. The method suggested by Huang and co-workers7 was used to calibrate the heat loss and to calculate the heat sink of hydrocarbon fuel. At each operating condition three measurements were performed to obtain the heat sink with reproducibility greater than 98.5%. The liquid product was drained into the residual tank, and the residual liquid sample was collected via a sampling valve in 1 min. Then it was weighed by an electrical balance (Mettler Toledo Co. Ltd.,) with an accuracy of 0.0001 g to obtain the residual flow rate. The gas yield (ψg) was obtained by the feed flow rate minus residual flow rates. Another method was also used to determine gas yields by the measured volume flow rate with a wet gas flowmeter, and the average molecular weight was analyzed using a micro GC 3000A (Agilent Inc., USA). The error between the two methods was less than 1.5%, indicating the reliability of the measurements. 2565


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Figure 2. Schematic diagrams of (a) experimental apparatus, (b) method, and (c) typical results for thermal cracking of hydrocarbon fuel. (a) I, feeding system; II, dc power system; III, electrically heated tube reactor; IV, data acquisition and controlling system; V, sampling and online analysis system. 1, feeding tank; 2, electronic balance; 3, filter; 4, high-pressure liquid chromatography pump; 5, mass flowmeter; 6, dc stabilized power supply; 7, safety valve; 8, pressure transducer; 9, copper bar; 10, data transmitting card, 11, electrically heated tube reactor; 12, condenser; 13, backpressure valve; 14, gas−liquid separator; 15, residual tank; 16, gas flowmeter; 17, data acquisition system. experimental errors and negligible terminal effect. Accordingly, it is reasonable to conclude that the developed methodology is indeed both practicable to experimentally obtain the local information on the cracked fuel and necessary to study the local heat transfer and coking behaviors of cracked fuel with the above information. 2.4. Analysis Methods. The gaseous products were analyzed online by a micro GC 3000A (Agilent Inc., USA), which has three thermal conductivity detectors (TCDs) and multichannel analytical columns, i.e., molecular sieve (10 m × 12 μm), Plot U (10 m × 30 μm), and alumina (10 m × 8 μm). Micro gas chromatography can detect hydrogen and other gaseous products within 1.5 min; all the gaseous products were identified and quantified by comparing with the

standard gas samples. The analysis deviation of gaseous products was less than 1.5%. More details about the analysis method can be found in ref 30. The liquid residual of the cracked fuel was identified by Agilent 5795 gas chromatography/mass spectrometry (GC−MS). An Agilent 7890A GC equipped with a flame ionization detector (FID) and an alkane−olefin−naphthene−aromatic (PONA) capillary column (50 m × 2 mm × 0.5 μm) was used to quantitatively analyze the content of each species in the liquid residuals. More details about the analysis method were reported previously by our group.30 In a typical run, more than 90% of the total fuel mass is quantified in the liquid 2566


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Figure 3. Representative chromatograms of liquid residuals from HF-I at different cracking conversions. Run 3: (a) HF-I fuel before cracking; (b) X = 39.9%; (c) X = 58.9%; (d) X = 75.1%; (e) X = 86.2%. products, and tens to hundreds of other species were found in negligible quantities (32) indicates the higher frequency of intermolecular Habstraction, especially at the lower conversions and temperatures. Aromatic hydrocarbons are the most important precursors of coke formation in the pyrolysis and soot formation in the fuel combustion. Figure 5f shows the aromatic yield along the tube. The major aromatics in the cracked fuel are present in the following order: toluene > benzene > ethylbenzene > styrene. They appear at the location of 55.5 cm corresponding to a conversion of 15%, indicating that aromatics may be secondary cracking products. Generally, aromatic yields increase monotonically with conversion and then steeply rise at high conversion (30%). Kumar and Kunzru37 observed aromatic yields of kerosene were less than 5% at 800 °C and residence time of 0.1 s. As we know, formation of aromatics is a typical secondary reaction, so the yields of aromatics should be lower for shorter residence time and lower reaction temperature. However, aromatic yields of 10% were observed in our work at 2569


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model was modified by introducing two lumped primary products of C5−C11 alkenes (C5+) and cycloalkenes (CC5+), as well as their secondary reactions. It was first assumed that kerosene was represented as a pseudopure compound and decomposed by a first-order reaction. The selectivities of primary products were assumed to be constant, and then the primary products underwent secondary reactions which were also represented by molecular reactions. The primary reaction can be represented as follows. kerosene → a1H 2 + b1CH4 + c1C2H4 + d1C2H6 + e1C3H6 + f1 C3H8 + g1C4 H8 + h1C4 H10 + i1C4 H6 + j1 C5 + + k1CC5 + + l1CnH 2n − 6

(4)

where a1, b1, ..., l1 are the stoichiometric coefficients of primary reactions. CnH2n−6 represents the heavy fractions, aromatics other than toluene, benzene, ethylbenzene, and styrene, and balancing carbon and hydrogen forming in primary reactions. The C5+ mainly consists of C5−C11 alkenes and trace alkanes, and CC5+ mainly includes cycloolefins and trace cycloalkanes. The average molecular weights for the C5+, CC5+, and CnH2n−6 were estimated to be 87, 85, and 120, respectively. The stoichiometric coefficients in eq 4 are determined from the initial selectivity of products defined as the number of moles of component formed per mole of raw material pyrolyzed. The selectivities of various products at different conversion levels were experimentally determined. There was no significant effect of operation conditions on initial selectivities. The selectivities of methane, hydrogen, ethane, and propane are approximately equal to constant values at conversions lower than 20% (Figure 7), indicating that a secondary reaction is insignificant at this region. The initial selectivities of ethylene and propylene could be represented as 0.5586 ± 0.0200 and 0.39 ± 0.0125, respectively. The yields of middle weight (C5−C11) alkenes change in similar trends; thus we plot yields of total middle weight alkenes as a lumped component. A similar treatment was also used for C5−C11 cycloalkenes. Table 4 lists the initial selectivities and coefficients for eq 4. The primary calculations using eq 4 and the secondary reactions of Kumar and Kunzru37 showed that the predicted yields of C5+ and CC5+ were higher than the experimental yields, whereas predicted yields of aromatics, including benzene, toluene, ethylbenzene, and styrene were lower than the experimental data. We noted that thermal cracking of alkenes underwent a similar pathway as that of alkanes, and that formation of aromatics was one of the most important pathways of cycloalkene pyrolysis.45,46 Therefore, the secondary reactions of C5+ and CC5+ were added into the reaction scheme.

Figure 6. (a) Calculated local densities and real linear velocities for run 4, (b) Reynolds number, and (c) residence time of cracked HF-I fuel along the tube. Solid line in (a), considering concentration changes of both gas and liquid phase; dashed−dotted line in (a), only gas phase concentration changes; dashed line in (a), without composition change. Lines 1, 2, 3, and 4 in (b) and (c) represent runs 1, 2, 3, and 4. Points A, B, C, and D in (c) represent beginning points of cracking.

which outlines the time scale of fluid flowing and cracking in the microchannel. Clearly, in all cases the residence times of fuel in the thermal cracking region are less than 0.05 s. 3.5. Molecular Kinetics Model. In general, there were three types of kinetics models for hydrocarbon cracking: detailed mechanism based, molecular kinetics, and lumped global models. The drawbacks of lumped global models in describing the secondary reactions also limited its potential applications. As mentioned before, it is also impractical to incorporate the thousands of elementary reactions in the engineering applications. The molecular kinetics model has been so far successfully used as a tool of predicting and optimizing the pyrolysis process of pure hydrocarbon, naphtha, and atmospheric gas oil mainly in the ethylene industry.37,43,44 Compared with the thermal cracking of naphtha feedstock, the appearance of middle weight (C5−C11) alkenes is obvious for the kerosene fuels. To account for this, the molecular kinetics

C5 + → a 2 H 2 + b2CH4 + c 2C2H4 + d 2C2H6 + e 2C3H6 + f2 C3H8 + g2C4 H8 + h2C4 H10 + i2C4 H6 CC5 + → a3B + b3T + c3 EB + d3ST + e3CnH 2n − 6

(5) (6)

where a2, b2, ..., i2 and a3, b3, ..., e3 are the stoichiometric coefficients of the secondary pyrolysis reactions of C5+ and CC5+, respectively. The stoichiometric coefficients were obtained as follows: first the C5+ and CC5+ fractions (boiling point lower than 150 °C, chromatography of C5+ and CC5+ is shown in Figure S-4 in the Supporting Information) were 2570


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Figure 7. Primary product selectivities as a function of conversion for different runs (■, run 1; ●, run 2; ▲, run 3; ▼, run 4). The selectivity of product i is defined as the moles of component i formed per mole cracked fuel. Yield of product i (also weight percent concentration) is defined as the mass percent of component i in bulk fuel.

Table 4. Initial Product Selectivities and Coefficients for eqs 4−6 (moles of Product Formed per mole Cracked)a

a

species

eq 4

H2 CH4 C2H4 C2H6 C3H6 C3H8 C4H8 C4H10 C4H6 C5+ CC5+ CnH2n−6 benzene toluene ethylbenzene styrene

a1 b1 c1 d1 e1 f1 g1 h1 i1 j1 k1 l1 − − − −

init selectivities for HF-I

init selectivities for HF-II

eqs 5 and 6

± ± ± ± ± ± ± ± ± ± ±

0.2219 0.7758 0.3865 0.3250 0.3961 0.1683 0.0368 0.1894 0.0093 0.2672 0.8228 0.0189 − − − −

a2 b2 c2 d2 e2 f2 g2 h2 i2 − −

0.1400 0.4800 0.3900 0.4500 0.0550 0.2700 0.3550 0.0955 0.0355 − −

a3 b3 c3 d3

0.7488 ± 0.0483 0.1396 ± 0.0136 0.05042 ± 0.00314 0.03402 ± 0.00122

0.1086 0.4773 0.5586 0.3900 0.4100 0.2001 0.2246 0.0353 0.0310 0.7201 0.2700 0.0222 − − − −

0.0102 0.0395 0.0200 0.0125 0.0324 0.0105 0.0098 0.0024 0.0020 0.0108 0.0230

secondary selectivities for HF-I and HF-II ± ± ± ± ± ± ± ± ±

0.0100 0.0204 0.0100 0.0200 0.0020 0.0189 0.0412 0.0037 0.0012

Conditions: 5 MPa, 25−705 °C (for HF-I fuel); 5 MPa, 25−715 °C (for HF-II fuel).

were determined through calculating the equivalent reactor volume.37,43 Pseudoisothermal is generally effective when there is no information available on the compositions and temperature along the tubular reactors. For each segment of an electrically heated tubular reactor defined as the length between two experimental points nearby, the averaged bulk temperature of inlet and outlet is taken as the reference temperature in this segment. Thus, a pseudoisothermal plug flow reactor was assumed for each segment. To build up a reactor model, some assumptions had to be made to reduce the complexity of the nonisothermal reactor, in which the composition and properties of the fuel changed site

separated from the cracked HF-I with conversions less than 20%, and then thermal cracking experiments were performed to obtain product distributions at a lower conversion less than 15%. The small molecular gaseous products were assigned as the products of C5+, while aromatics were assigned as the products of CC5+. Finally, the stoichiometric coefficients were determined following similar approaches as described above. 3.6. Reactor Model of Electrically Heated Tube. Previously, a pseudoisothermal approach of a tubular reactor was usually utilized that the measured temperature profile was converted to isothermal conditions at a chosen reference temperature and pressure, and the global kinetic parameters 2571


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Table 5. Molecular Reaction Kinetics Scheme for Thermal Cracking of HF-I Fuela Ea, kJ mol−1

reactionb

k0, s−1

source

HF‐I (C11.85H 23.82) → 0.1086H 2 + 0.4773CH4 + 0.5586C2H4 + 0.39C2H6 + 0.41C3H6 + 0.2001C3H8 + 0.2246C4 H8 + 0.0353C4 H10 + 0.031C4 H6 + 0.7201C5 + + 0.27CC5 + + CnH 2n − 6

C2H6 ↔ C2H4 + H 2

2.869 × 1014

this work

272.6

4.652 × 1013

36

273.1

7.284 × 1012

36

(R1)

(R2)

C3H6 ↔ C2H 2 + CH4

217.9

(R3)

C2H 2 + C2H4 → C4 H6

(R4)

172.5

(1.026 × 10 )

36

2C2H6 → C3H8 + CH4

(R5)

272.8

3.75 × 1012

36

252.6

(7.083 × 10 )

36

189.4

5.0 × 1012

36d

211.5

4.692 × 10

246.9

(2.536 × 10 )

36

244.9

1.200 × 1012

36

228.1

1.424 × 10

36d

250.8

(1.0 × 10 )

36

190.3

7.8 × 10

36d

295.4

7.0 × 10

C2H4 + C2H6 → C3H6 + CH4

C3H8 ↔ C3H6 + H 2

(R6)

(R7)

C3H8 → C2H4 + CH4

(R8)

C3H8 + C2H4 → C2H6 + C3H6

2C3H6 → 3C2H4

(R9)

(R10)

2C3H6 → 0.3CnH 2n − 6 + 0.14C5 + + 3CH4 C3H6 + C2H4 → C4 H8 + CH4 n ‐ C4 H10 → C3H6 + CH4

n ‐ C4 H10 → 2C2H4 + H 2

(R12)

(R13) (R14)

n ‐ C4 H10 → C2H4 + C2H6

n ‐ C4 H10 → C4 H8 + H 2

(R11)

(R15) (R16)

1 ‐ C4 H8 → 0.41CnH 2n − 6 + 0.19C5 + 1 ‐ C4 H8 ↔ C4 H6 + H 2

(R17)

(R18)

9 c

10 c

10 10 c

11

11 c

12 14

36

36

256.3

4.099 × 10

12

36

260.7

1.637 × 1012

36

195.2

1.075 × 10

36d

209.0

1.0 × 1010

36d

13

C4 H6 + C2H4 → B + 2H 2

(R19)

231.0

(2.774 × 1013)c

36d

C4 H6 + C3H6 → T + 2H 2

(R20)

240.6

(1.720 × 1014)c

36d

193.9

(1.000 × 10 )

36d

181.2

4.0 × 1010

36d

189.6

1.231 × 1013

this work

194.4

9.6935 × 1012

this work

C4 H6 + 1 ‐ C4 H8 → EB + H 2

2C4 H6 → ST + 2H 2

(R21)

(R22)

11 c

C5 + → 0.14H 2 + 0.48CH4 + 0.39C2H4 + 0.45C2H6 + 0.055C3H6 + 0.27C3H8 + 0.355C4 H8 + 0.0955C4 H10 + 0.0355C4 H6 + 0.1091CnH 2n − 6

(R23)

CC5 + → 0.7488B + 0.1396T + 0.05043EB + 0.03402ST + 0.04262Cn H 2n − 6

(R24)

Conditions: 5 MPa, 25−705 °C (for HF-I fuel); 5 MPa, 25−715 °C (for HF-II fuel). bB, benzene; T, toluene; EB, ethylbenzene; ST, styrene. Units: m3·mol−1·s−1. dKinetics parameters regressed by authors.

a c

by site. In this work, the cooling channel was modeled as a onedimensional plug flow reactor whose axial temperature profile was defined by a fourth-order polynomial of the experimental values. In a plug flow reactor, the steady-state continuity equation for component i can be written as dFi = S dl

where CA, j =

∑ (vrirn) n=1

dl

⎛ −E ⎞ = k 0 exp⎜⎜ a ⎟⎟CA, j ⎝ RTj ⎠

P ZjRTj

(9) (10)

ν and ε are calculated from the experimental product distribution for each run, and CA is the concentration of HF-I kerosene in the reactor. The compressibility factor Zj of cracked kerosene is calculated from the Peng−Robinson equation of state using NIST SUPERTRAPP software. 3.7. Kinetics Constants of Primary and Secondary Reactions. Kumar and Kunzuru37 developed first-order kinetics of the primary pyrolysis of naphtha and obtained the reaction rate constant involving an expansion factor with an equivalent reactor volume and pseudoisothermal approach. Following this approach, eq 11 was derived for each segment of an electrically heated tube.

(7)

where N represents the number of reactions involving component i in Table 5. With reaction kinetic constants presented in Table 5, eq 7 for 18 components was numerically integrated using a fourth-order Runge−Kutta−Gill algorithm. For the aviation fuel (in segment j), the equations above can be represented as

S

1 + εXA, j

εj = νj − 1

N

FA, j dXA, j

1 − XA, j

(8) 2572


Energy & Fuels ⎛ 1 − X1, j ⎞ ⎟⎟ − εj(X 2, j − X1, j) kjt j = (1 + εj) ln⎜⎜ ⎝ 1 − X 2, j ⎠

Article

benzene with only slight overpredictions. Figure 10a further compares the experimental and calculated yields for all species. It is observed that more than 98% points fall within the deviations of ±20% with an average absolute relative error (AARE) equal to 9.2%. Clearly, the deviations of some points (especially when the yields are less than 4%) may be caused by the following facts: the possible experimental and analysis errors for the complex compositions of hydrocarbon fuels; the limited components and reactions used in the developed kinetics model; the possible changes in the chemical compositions of middle weight fractions under higher conversions; the preliminary assumptions of plug flow reactor models. The value of F = 2492.49 for the F-test of the kinetics model was higher than Fcrit = 1.54, confirming that the regressed model fitted well the observed values. The sensitivity of the model was checked by changing the rates of each secondary reaction by 20%. It was found that only the rates of reactions R10, R11,R17, R23, and R24 in Table 5 led to a minor change in the product yields when the rates of the other secondary reactions were increased. The experimental and calculated heat sinks of hydrocarbon fuels are compared to show the consistency and validity of the developed models. The experimental heat sinks are measured by the total heat input (I2R) minus the heat lost (calibrated by the correlation of the empty tube temperature and the heat power), according to the method proposed by Huang et al.7 The predicted heat sinks were calculated from the enthalpy difference between the inlet and calculated outlet fuel compositions and temperatures using the SUPERTRAPP software. Figure 10b plots the heat sink of HF-I as a function of bulk fuel temperature. The physical heat sink of HF-I fuel almost linearly increases up to 1500 kJ/kg when the fuel temperature rises from room temperature to 550 °C. After that the heat sink of the fuel sharply increases to 3045 kJ/kg at 680 °C where the thermal cracking conversion reaches up to 86.2%. The turning point at 550 °C is in good consistency with the beginning point of the thermal cracking of run 4.The chemical heat sink attributed to thermal cracking reactions sharply contributes >33% of the total heat sink (ca. 1000 kJ/kg). In addition, the predicted heat sinks agree well with the experimental values, indicating that the thermal cracking of supercritical HF-I fuel is well described with the developed models, even in thermal cracking conversions as high as 86%. 3.9. Thermal Cracking of HF-II Fuel. To further verify the validation of the developed experimental method as well as the kinetics model for hydrocarbon fuels, thermal cracking of HF-II fuel (run 5) was used following the experimental and kinetic approaches described above. It should be noted that for HF-II both the coefficients in the primary reactions (as shown in Table 4) and kinetics parameters (Ea of 207.6 kJ·mol−1, k0 of 2.4351 × 1014 s−1) were experimentally determined following the method described above. Obviously, the predicted yields of HF-II as well as its gaseous and liquid products are in good agreement with experimental results (see Figure 11). Consequently, the developed kinetics model could be used to reliably describe the thermal cracking of any other hydrocarbon aviation fuel through a series of experimental runs following the method proposed in this work.

(11)

where j represents the jth segment of the reaction tube, 1 and 2 represent the inlet and outlet of segment j. Residence time t in each segment is calculated from the inlet and outlet average velocity. The cracking rate of kerosene for each tube segment is calculated from eq 11. Figure 8 shows the Arrhenius plot for

Figure 8. Arrhenius plot for the primary thermal cracking of supercritical HF-I fuel. ■, Run 1; ●, run 2; ▲, run 3; ▼, run 4.

the thermal decomposition of kerosene. The overall decomposition of kerosene was represented by first-order kinetics with Ea of 217.9 kJ·mol−1 and k0 of 2.8686 × 1014 s−1, which are in good agreement with previous results reported for kerosene jet fuel. For instance, Widegren et al.26,47 reported Ea of 220 kJ·mol−1 for the thermal decomposition of Jet-A jet fuel, and Ea of 201 kJ·mol−1 for RP-1 and RP-2 fuels. The set of secondary reactions proposed for naphtha pyrolysis was first tried, and the kinetic parameters of several secondary reactions were adjusted by trial and error to minimize the deviation between the experimental and predicted yields. The final values of the pre-exponential factor and activation energy for each reaction are presented in Table 5. Compared with the original kinetic parameters presented in ref 37, the kinetic parameters of several reactions (as shown in Table 5) are modified to satisfactorily simulate the variation of yields of propylene and aromatic hydrocarbons for each run. Additional reactions R23 and R24 are presented for the consumption reactions of C5+ and CC5+, and their kinetic constants were regressed from the experimental data. We must say that for HF-I fuel all model predictions were determined using the four experimental runs as controls, while for HF-II all the equations in Table 5 were used, but the coefficients for reaction R1 were experimentally determined. 3.8. Comparisons of Predicted and Experimental Results. Figure 9 compares the experimental yields of both gaseous and liquid products and predicted values from the developed model in this work. For all the experimental runs, predicted yields for most of the major components are generally in agreement with the experimental values. For instance, the calculated yields of methane, ethane, ethylene, and propylene are in good agreement with the experimental results over a wide range. The yields of C5+ and CC5+ are also well predicted by the model with a good accuracy within ±3%. For the aromatics results, satisfactory agreement with experimental results is achieved for the predicted yields of toluene and

4. CONCLUSIONS In this work, a new experimental methodology of an electrically heated tube offered new insights into the thermal cracking of EHFs in regenerative cooling microchannels. This method 2573


Energy & Fuels

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Figure 9. Comparison of predicted and experimental yields of typical products in different runs. ■, Run 1; ●, run 2; predicted values.

characteristically involved cracking reactions carried out at variable tube length with constant electrical current. For the first time the local chemical compositions of cracked hydrocarbon fuels downstream along the cooling microchannel under four typical conditions (5 MPa, >80% conversion) were reported in detail. It was found that for the kerosene-type hydrocarbon fuel the middle weight 1-alkenes and cycloalkenes account for a large amount of primary decomposition products. To describe the thermal cracking of hydrocarbon fuels in a cooling channel, a molecular reaction scheme of modified Kumar−Kunzru kinetics model consisting of 18 species and 24 reactions (1 primary reaction and 23 secondary reactions) was proposed with introducing middle weight alkenes as two

▲,

run 3;

▼,

run 4; lines,

dependent lumped components. With the help of a simplified reactor model, the kinetics parameters of both primary and secondary reactions were obtained. The reliability of the developed model was confirmed through comparing the experimental compositions and heat sinks with the corresponding predicted values, as well as the heat sink curve. It was concluded that the proposed kinetics model was able to predict the thermal cracking behaviors of hydrocarbon fuels as well as the detail chemical compositions of cracked jet fuel even at very high conversions up to 86%. The methodology proposed in this work could be easily extended to the thermal cracking of any hydrocarbon fuel. 2574


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Article

ASSOCIATED CONTENT

S Supporting Information *

Some experimental and calculation results of Table S-1, and Figures S-1−S-4. This material is available free of charge via the Internet at http://pubs.acs.org.

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AUTHOR INFORMATION

Corresponding Author

*Tel./fax: +86-22-27892340. E-mail: [email protected] (Guozhu Liu). Author Contributions †

Both authors contributed to this work equally.

Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS The authors gratefully acknowledge financial support from the National Natural Science Foundation of China (Grant 91116001) and the Programme of Introducing Talents of Discipline to Universities (B06006). Grateful thanks are also given to Dr. Fanxu Meng of Texas A&M University and Prof. Li Wang and Prof. Zhentao Mi of Tianjin University for their careful proofreading, and to Miss Haijin Li for her kind help in the composition analysis of pyrolysis products. G.L sincerely expresses his special thanks to a previous reviewer of JFUE-D12-01654 for constructive suggestions.

Figure 10. (a) Parity plot of experimental and predicted values and (b) heat sink of HF-I as a function of bulk temperatures. Points, experimental results of this work; line, predicted results with developed model in this work.

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NOTATION Ai = chromatographic peak area of i C = concentration, mol/m3

Figure 11. Thermal cracking of HF-II downstream distribution along electrically heated tube (run 5). (a) Fuel temperature and conversion; (b) major gas products; (c) C5+ and CC5+; (d) toluene and benzene. Lines calculated results; points, experimental results. 2575


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Advanced Distillation Curve Method. Energy Fuels 2009, 23 (7), 3637−3644. (15) Linne, D. L.; Meyer, M. L.; Edwards, T.; Eitman, D. A. Evaluation of Heat Transfer and Thermal Stability of Supercritical JP-7 Fuel. Proceedings of the 33rd Joint Propulsion Conference and Exhibit; American Institute of Aeronautics and Astronautics: Reston, VA, 1997; DOI: 10.2514/6.1997-3041. (16) Zhong, F. Q.; Fan, X. J.; Yu, G.; Li, J. G.; Sung, C. J. Heat Transfer of Aviation Kerosene at Supercritical Conditions. J. Thermophys. Heat Transfer 2009, 23 (3), 543−550. (17) Guo, W.; Zhang, X. W.; Liu, G. Z.; Wang, J.; Zhao, J.; Mi, Z. T. Roles of hydrogen donors and organic selenides in inhibiting solid deposits from thermal stressing of n-dodecane and Chinese RP-3 jet fuel. Ind. Eng. Chem. Res. 2009, 48 (18), 8320−8327. (18) Yang, C.; Liu, G.; Wang, X.; Jiang, R.; Wang, L.; Zhang, X. Preparation and anti-coking performance of MOCVD alumina coatings for thermal cracking of hydrocarbon fuels under supercritical conditions. Ind. Eng. Chem. Res. 2012, 51 (3), 1256−1263. (19) Sheu, J. C.; Zhou, N.; Krishnan, A.; Jones, A. G.; Katta, V. R. Modeling of pyrolysis and pyrolytic deposition of Norpar-13 under supercritical conditions. Prepr.Am. Chem. Soc., Div. Pet. Chem. 1998, 43 (3), 390−393. (20) Sheu, J. C.; Zhou, N.; Krishnan, A.; Jones, A. G.; Katta, V. R. Thermal cracking of Norpar-13 under near-critical and supercritical conditions. AIAA Pap. 1998, 3758, 629. (21) Ward, T. A.; Ervin, J. S.; Striebich, R. C.; Zabarnick, S. Simulations of flowing mildly-cracked normal alkanes incorporating proportional product distributions. J. Propul. Power 2004, 20 (3), 394− 402. (22) Ward, T. A.; Ervin, J. S.; Zabarnick, S.; Shafer, L. Pressure effects on flowing mildly-cracked n-decane. J. Propul. Power 2005, 21 (2), 344−355. (23) Bao, W.; Li, X.; Qin, J.; Zhou, W.; Yu, D. Efficient utilization of heat sink of hydrocarbon fuel for regeneratively cooled scramjet. Appl. Therm. Eng. 2011, 33−34, 208−218. (24) Hou, L.; Dong, N.; Sun, D. Heat transfer and thermal cracking behavior of hydrocarbon fuel. Fuel 2013, 103, 1132−1137. (25) Zhong, F.; Fan, X.; Yu, G.; Li, J. Thermal cracking of aviation kerosene for scramjet applications. Sci. China, Ser. E 2009, 52 (9), 2644−2652. (26) Widegren, J. A.; Bruno, T. J. Thermal decomposition kinetics of kerosene-based rocket propellants. 1. Comparison of RP-1 and RP-2. Energy Fuels 2009, 23 (11), 5517−5522. (27) Dahm, K. D.; Virk, P. S.; Bounaceur, R.; Battin-Leclerc, F.; Marquaire, P. M.; Fournet, R.; Daniau, E.; Bouchez, M. Experimental and modelling investigation of the thermal decomposition of ndodecane. J. Anal. Appl. Pyrolysis 2004, 71 (2), 865−881. (28) Herbinet, O.; Marquaire, P. M.; Battin-Leclerc, F.; Fournet, R. Thermal decomposition of n-dodecane: experiments and kinetic modeling. J. Anal. Appl. Pyrolysis 2007, 78 (2), 419−429. (29) Li, X. F.; Huai, X. L.; Cai, J.; Zhong, F. Q.; Fan, X. J.; Guo, Z. X. Convective heat transfer characteristics of China RP-3 aviation kerosene at supercritical pressure. Appl. Therm. Eng. 2011, 31 (14− 15), 2360−2366. (30) Jiang, R.; Liu, G.; He, X.; Yang, C.; Wang, L.; Zhang, X. Supercritical thermal decompositions of normal- and iso-dodecane in tubular reactor. J. Anal. Appl. Pyrolysis 2011, 92 (2), 292−306. (31) Van Camp, C. E.; Van Damme, P. S.; Froment, G. F. Thermal cracking of kerosene. Ind. Eng. Chem. Process Des. Dev. 1984, 23 (1), 155−162. (32) Van Camp, C. E.; Van Damme, P. S.; Willems, P. A.; Clymans, P. J.; Froment, G. F. Severity in the pyrolysis of petroleum fractions. Fundamentals and industrial application. Ind. Eng. Chem. Process Des. Dev. 1985, 24 (3), 561−570. (33) Fukiba, K.; Tsuboi, N.; Minato, R. Numerical study on the heat transfer of the flow with endothermic chemical reaction. 45th AIAA/ ASME/SAE/ASEE Joint Propulsion Conference and Exhibit; American Institute of Aeronautics and Astronautics: Reston, VA, 2009; DOI: 10.2514/6.2009-5461.

Ea = activation energy, kJ/mol FA0, Fi = molar flow rates of inlet kerosene and component i, mol/s f i = correction factor g = heat loss function h = heat flux, kW/m2 ΔH = heat sink, kJ/kg I = electric current, A k = decomposition kinetic constant, L/(mol·s) k0 = pre-exponential factor, 1/s or m3·mol−1·s−1 L = heating length, m m = mass flow rate, g/s P = pressure, MPa Qin, Qloss = input heat and loss heat, kJ R = electricity resistance, Ω S = cross-sectional area of the tube, m2 T = temperature, K t = residence time, s X = conversion, % X1, X2 = inlet and outlet cracking conversions of each segment Z = compressibility factor Greek Symbols

ε = expansion coefficient κ = volume resistivity, Ω·m ν = moles of product formed per mole of kerosene decomposed ψg = gas yield

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