Experimental study of pyrolysis-combustion coupling in a regeneratively cooled combustor: Heat transfer and coke formation

Fuel 239 (2019) 1091–1101

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Experimental study of pyrolysis-combustion coupling in a regeneratively cooled combustor: Heat transfer and coke formation

T

L. Taddeoa, , N. Gascoina, K. Chetehounaa, A. Ingenitob, F. Stellab, M. Bouchezc, B. Le Naourc ⁎

a

INSA Centre Val de Loire, 88 Boulevard Lahitolle, Bourges, France University of Rome “La Sapienza”, Department of Mechanics and Aeronautics, 18 Via Eudossiana, Rome, Italy c MBDA France, 8 rue Le Brix, Bourges, France b

ARTICLE INFO

ABSTRACT

Keywords: Regenerative cooling Hydrocarbon pyrolysis Combustion-pyrolysis coupling Hydrocarbons coking activity Heat transfer

Scramjets engines are suitable to propel high-speed hypersonic vehicles. As flight velocities increase, vehicle thermal protection becomes very critical. In this sense, regenerative cooling is a well-known cooling technique, particularly effective when an endothermic hydrocarbon is used as fuel. The development of regeneratively cooled engines faces several challenges, the most important being the difficulty of defining an engine regulation strategy because of the dual function of the fuel (both propellant and coolant). In this context, a regeneratively cooled combustor allowing the experimental study of a fuel-cooled engine has been designed. Experiments are run using ethylene as fuel and air as oxidizer. Two command parameters, i.e. fuel mass flow rate and equivalence ratio (1.0–1.5), are investigated. It has been observed that fuel mass flow rate increases by 16–20% result in heat flux density (from the combustion gases to the combustor wall) increases between 20 and 28%, depending on equivalence ratio and pressure. The dependence of the cooling system heat exchange efficiency on the two operating parameters has been demonstrated. Ethylene coking activity has been investigated. For applied interest, a monitoring method for carbon deposits formation has been developed and validated.

1. Introduction Scramjet is an air-breathing jet engine, designed to trust vehicles flying in the atmosphere at a very high speed (at least Mach 4). It is a very simple motor, with no moving parts. Air rushes in the air inlet at very high speed and is compressed thanks to the forward motion of the vehicle, without any rotary compressor. According to several studies, Scramjet could permit to achieve hypersonic flight, above Mach 5 [1,2]. Hypersonic flight fits to various technological applications, as air launched strike weapons, very high speed aircrafts, reusable space transport vehicles [3]. However, as flight velocity increases beyond the supersonic one, the difficulties accompanying aircraft design increase, notably with respect to the combustion chamber, which is exposed to very high temperatures (combustion gases total adiabatic temperature could even achieve 4500 K) [4–8]. To maintain the combustor structures at a low and safe temperature, several cooling techniques can be used. In general, they can be divided into two categories: passive ones, where no coolant is used, and active ones, where a coolant is used [7,9,10]. Active cooling techniques consist in the use of a coolant that passes through cooling channels located in the combustion chamber walls. In this sense, two main approaches

⁎

can be considered, as it is possible to use as coolant a fraction of the air entering the air inlet (active cooling by air) or even the fuel (active cooling by fuel). When compared to air, fuels are generally much better coolants. Hence, active cooling by fuel is usually considered the best cooling strategy and the only adequate to deal with the very high temperatures and the very high heat fluxes characterizing Scramjet engines [11]. Regenerative cooling is one of the most well-known active cooling techniques. The fuel, acting as the coolant, flows through cooling channels which are located between the inner and the outer walls of the engine, before being burned. A counter-flow heat exchange between the fuel-coolant and the combustion gases is thus generated [12,13]. This technology is particularly effective when an endothermic hydrocarbon is used as propellant, because of two main reasons: i) when heated at high temperatures, endothermic hydrocarbons undergo thermal pyrolysis and the endothermic behavior of the decomposition reactions enhances their cooling capacity [7,12]; ii) fuel decomposition generates light species (mainly hydrogen, ethane and ethylene) whose ignition delay times are very low when compared to that of the propellant. The latter is a key benefit because one of the weak points of Scramjets engines, deriving from the very low time that the fuel and the oxidizer

Corresponding author at: 88 Boulevard Lahitolle, Bourges 18000, France. E-mail addresses: [email protected] (L. Taddeo), [email protected] (N. Gascoin).

https://doi.org/10.1016/j.fuel.2018.11.096 Received 15 May 2018; Received in revised form 14 November 2018; Accepted 20 November 2018 0016-2361/ © 2018 Published by Elsevier Ltd.

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Nomenclature A Cp D ΔP F h H k L ṁ P q S t T v

η φ μ ρ σ τ

surface, m2 heat capacity, J.kg−1.K−1 diameter, m pressure drop, Bar view factor, dimensionless convective heat transfer coefficient, W.m−2.K−1 height, m thermal conductivity, W.m−1.K−1 length, m mass flow rate, kg.s−1 pressure, Bar heat flux density, W.m−2 cross section, m2 time, s temperature, K velocity, m.s−1

Subscripts amb cc conv comb f g hf w,int rad th

Greek symbols ε

thermal exchange efficiency, dimensionless equivalence ratio of fuel to oxidizer, dimensionless dynamic viscosity, Pa.s density, kg.m−3 Boltzmann’s constant, W.m−2.K−4 Fuel pyrolysis rate, dimensionless

ambient conditions cooling channel convective combustor fuel fed to the cooling channel combustion gases fuel at cooling channel outlet, before injection in the burner combustor internal wall radiative thermocouple

emissivity, dimensionless

spend in the combustor (in the order of few ms [13]), is the difficulty to complete fuel combustion in the combustion chamber. Consequently, fuel pyrolysis improves the engine performance by enhancing fuel combustion. The most important issues concerning the design and development of regeneratively cooled Scramjets have already been studied in the framework of various programs. The evaluation of the heat flux density between the combustion gases and the combustor internal surfaces, the identification of the highest heat flux regions, the optimization of the cooling channels shape and length, etc., have been the subjects of a number of studies [5,14–25]. Several studies have also been realized on endothermic hydrocarbons thermal decomposition, aiming at determining the main products for various jet fuels and operating conditions of pressure, temperature or residence time [4,26–29]. Because of its possible harmful effects on the efficiency of cooling system, a particular attention has

been paid to the coking activity resulting from their high-temperature pyrolysis. Indeed, especially at very high temperatures (above 1400 K), hydrocarbons decomposition mainly produces species as hydrogen, ethylene and ethane, which are, when compared to the most common jet fuels, highly hydrogenated. Thus, carbon depositions are formed [6,7,30–34]. This represents a very important challenge facing the use of high-temperature decomposing hydrocarbons on regeneratively cooled vehicles, especially where very long system lives are required. Hydrocarbons coking activity has at least three negative consequences: i) decrease in the heat transfer efficiency of the overall cooling system (due to thermal insulation of carbon, whose thermal conductivity is 4–6 times less than that of steel metallic materials) [30]; ii) decrease in the endothermicity of fuel decomposition reactions [7]; iii) increase in the pressure drop (or even system failure) due to cooling channels or fuel injectors obstruction [31]. Nevertheless, to the author’s knowledge, none of the above

Fig. 1. Scheme of the regeneratively cooled combustion chamber. 1092

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– fuel pyrolysis coupling in a regeneratively cooled engine has been designed. Thanks to it, the effects of fuel mass flow rate variations on some of the main output parameters, as fuel temperature, pressure and residence time in the cooling channels, flame temperature, etc., have been studied. The impact of another operating parameter, i.e. fuel to oxidizer equivalence ratio, has also been investigated. To refine the understanding of the results of the experiments, numerical calculations have been realized using the software CHEMKIN-PRO to model fuel pyrolysis before injection in the combustor. Some results of this study have already been presented in terms of the influence of the two operating parameters on the dynamics of the combustor [35]. The present article focuses more specifically on the impact of these parameters on the convective and the radiative heat fluxes passing from the burned gases to the combustor wall and the heat flux absorbed by the fuel-coolant in the cooling channels. The heat transfer efficiency of the cooling system in terms of energy recovery is determined. The coking activity of the fuel is analysed; a carbon deposition monitoring method, suitable for real-time on-board application, is proposed and validated. This work, along with the ones already published, will raise the knowledge of the scientific community on the thermal management of a regeneratively cooled Scramjet. It will allow developing analytical relationships between the operating parameters and the measured outputs, permitting to define an engine control and regulation strategy in future works.

Table 1 Test cases for different operating conditions. Test case

Fuel mass flow rate · 105 (kg s−1)

Equivalence ratio

Maximum fuel pressure during the test (Bar)

1 2 3 4 5 6 7 8 9

2.0–3.0–4.0 2.0–2.4–2.8 2.0–2.4–2.8 2.0–2.4–2.8 2.4 2.0–2.4–2.8 2.0–2.5–2.9 2.0–2.4–2.9 1.6

1.50 1.00 1.25 1.50 1.00–1.25–1.50 1.00 1.25 1.50 1.25

13.8 3.8 3.2 4.0 7.2 6.7 11.0 32.2 10.5

mentioned studies focused on the following key application issue, which is the need to define a proper engine control strategy. Due to the strong coupling between fuel decomposition and fuel combustion, resulting from the dual function of the propellant, the effect on the engine thrust of the most important operating parameter, i.e. the mass flow rate of the fuel fed to the combustion chamber, is not easy to determine. In fact, because of the complexity of the physical and chemical phenomena involved (heat transfer between the channels and the combustion chamber, fuel pyrolysis, fuel combustion, etc.) and to their reciprocal interactions, an increase in fuel mass flow rate, which should lead to an increase in the engine thrust, could result in a decrease in this thrust or even in flame extinction [12,13]. To improve understanding of the interaction between fuel pyrolysis and combustion on the behavior of the engine when the operating parameter is varied, a comprehensive study, both experimental and numerical, is necessary. To this end, an innovative experimental set-up, consisting of a labscale fuel-cooled combustor suitable for the analysis of fuel combustion

2. Materials and methods 2.1. Description of the experimental set-up The experimental set-up has been described previously [35]. The

Fig. 2. Sequence representing how thermocouples are moved during each test. 1093

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combustion chamber, whose scheme is given in Fig. 1, has an internal a diameter of 0.1 m. Before being fed to the burner, the fuel flows through a 45 m long rolled-up stainless steel tube (indicated in Fig. 1 as “cooling channel”) that passes in the combustion chamber, entering from the top and exiting from the bottom. Due to this configuration, corresponding to a counter-flow heat exchanger, the fuel, exposed to the heat flux generated by the flame, decomposes before being burned. Tests are run using ethylene as fuel and air as oxidizer. The choice of ethylene as fuel is justified by the fact that this combustible is one of the main products of kerosene and of other jet fuels thermal decomposition. Hence, the use of this hydrocarbon permits to characterize the behavior of a fuel-cooled engine under more realistic operating conditions [28,29]. The two gases are fed to the combustor by using two highprecision mass flow controllers, respectively a Bronkhorst F-201CM10 K-RDA-88-K for ethylene (with a range of 0–2·10−4 kg.s−1) [36] and a Bronkhorst F-202AV-M20-RDA-55-V for air (with a range of 0–5·10−3 kg.s−1) [37]. According to the manufacturer, the accuracy of both mass flow controllers is ± 2% of the full scale. Two pressure transducers, indicated as PTE-1 and PTE-2 in Fig. 1, are used to measure ethylene pressure respectively at the inlet of the cooling channel and at the inlet of the burner. Another pressure transducer, indicated as PTA in Fig. 1, measures the pressure of the air entering the burner. Due to security reasons, the authors decided to carry out fuel combustion at ambient pressure. Nevertheless, to be closer to scramjet real operating conditions, a specific nozzle permitting to increase the pressure of the fuel flowing in the rolled-up cooling channel was designed and installed at cooling channel outlet, before pressure transducer PTE-2. This device, indicated as PC (Permeation Cell) in Fig. 1, has been described in a previous work, where its operating principles have been comprehensively detailed [13]. The temperatures achieved by the burned gases are measured by eleven thermocouples, which can be shifted from the wall to the axis of the combustor. A thermocouple (indicated as TChf in Fig. 1) is used to measure the temperature of the heated fuel.

absorbed by the fuel-coolant in the cooling channel and heat transfer efficiency of the cooling system in terms of energy recovery. Heat transfer from the burned gases to the combustor internal surface occurs both by convection and by radiation:

q g = q g,conv + q g,rad

(1)

In the above equation qg is the total heat flux from the burned gases to the combustor internal surface, qg,conv is the convective one and qg,rad is the radiative one. The following equation is used to calculate qg,conv:

q g,conv = hg (Tg,P3

Tg,P1)

(2)

where hg is the convective heat transfer coefficient, Tg,P1 is the gas temperature measured in position P1, and Tg,P3 is the gas temperature measured in position P3. The convective heat transfer coefficient hg is calculated as [38]:

kg

hg = 1.86

Dw,int Reg Prg H

Di

0.33

µg,P3

0.14

µg,P1

(3)

where Reg and Prg are defined as follows:

Prg =

Dw,int vg

g

Reg =

µg

(4)

C p,g µg (5)

kg

In Eqs. (3)–(5), kg is the average gas thermal conductivity, ρg is the average gas density, μg is the average gas viscosity, Cp,g is the average gas specific heat, vg is the average gas velocity, Dw,int is the internal diameter of the combustor, H* is a characteristic length, μg,P1 is gas viscosity in position P1, and μg,P3 is gas viscosity in position P3. The average values given above are calculated at the average gas temperature between positions P1, P2, and P3. The following equation is used to calculate qg,rad [39]:

q g,rad =

2.2. Test methodology

0

F (Tg,P3 4

Tg,P14)

(6)

where

A brief review of the tests that have been run is proposed in Table 1. Every experiment is carried out by varying a single input parameter, whether fuel mass flow rate or equivalence ratio, in order to analyze its impact on the main outlet parameters under steady and transient conditions. During each test, this parameter is increased twice and then decreased twice, while keeping the other constant. This enables catching the hysteresis phenomena characterizing regeneratively cooled combustion chambers whose presence has been detailed in previous works [12,13]. During a test, operating conditions are varied only when steady state has been achieved; thus, during each experiment, the test bench achieves steady state five times. For simplicity purposes, these steady states will be numbered consecutively, beginning from steady state 1 and finishing with steady state 5. As it can be seen in Fig. 2, at the start of each test all the thermocouples are placed at 1 cm from the wall of the combustor, at position P1, with the exception of the two thermocouples indicated respectively as TC2 and TC7, which are placed at 5 cm from the wall, on the axis of the combustion chamber, at position P3 (Fig. 2-a). When a steady state is achieved, the thermocouples which are at position P1 are first shifted at position P2 (at 3 cm from the wall, Fig. 2-b), then at position P3 after the stabilization of the measured temperatures (Fig. 2-c). Before varying the input parameter under study, the thermocouples which had been displaced are moved back to their original positions (Fig. 2-d). Further details on the test methodology can be found in [35].

0

=

1

+

g

1

1 (7)

p

In the above equations, εg is the combustion gases emissivity, εp is the combustor wall emissivity, σ is the Stefan–Boltzmann’s constant, and F is the view factor. Combustion gases emissivity εg was calculated using an empirical correlation proposed by Leckner [40]: g

=

CO

+

CO2

+

H2 O

CO2 H2O

(8)

where εCO, εCO2 and εH2O are respectively carbon monoxide, carbon dioxide and water vapor emissivity and ΔεCO2 is a correction factor. The heat flux absorbed by the fuel in the cooling channel is given by the following equation: chem qf,cc = qsens f,cc + qf,cc

(9)

where and represent respectively the sensible and the chemical heat flux absorbed by the pyrolyzing fuel (the latter because of the endothermicity of pyrolysis reactions). The following equation is used to calculate the sensible heat flux qfsens , cc :

qfsens , cc

qfchem , cc

qsens f,cc = mf C p,f (Thf

Tf )

(10)

where ṁf is fuel mass flow rate, Cp,f is the average fuel specific heat in the cooling channel, Tf is fuel temperature at cooling channel inlet, and Thf is fuel temperature at cooling channel outlet, before its injection in the burner. Fuel specific heat is a function of fuel composition and temperature, which vary along the cooling channel following ethylene heating and

2.3. Data post-processing The objectives of this work include the calculation of the heat load passing from the combustion gases to the combustor wall, the heat load 1094

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pyrolysis. The evolution of the physical and chemical properties of the fuel along this channel has been determined by means of numerical calculations realized with the Plug Flow Reactor module of the CHEMKIN-PRO package [41,42]. In order to model ethylene decomposition, a chemical reaction mechanism developed by the DCPR laboratory in Nancy to simulate n-dodecane thermal decomposition has been used [43]. Although this reaction scheme is not specifically dedicated to the thermal decomposition of ethylene, it permits to model the pyrolysis of this hydrocarbon. In fact, because of its highly developed mechanism, considering 153 molecular species, 118 radical species and 1185 chemical reactions [12], it can be used to take into account not only the degradation of n-dodecane (primary scheme) but also that of the main by-products of its decomposition. Indeed, since its development was aimed at predicting the formation of carbon deposits, which follows n–dodecane decomposition and is highly dependent on the by-products generated by it, it also allows to model the thermal decomposition of ethylene, which is one of main by-products of n-dodecane pyrolysis [43]. Further details on these CHEMKIN-PRO numerical calculations can be found in [44]. As it will be explained in Section 3, the maximum temperature reached by the fuel in the cooling channel is probably higher than the one measured at its exit, Thf. Indeed, after having been exposed to the heat flux generated by the flame, this channel passes through a noninsulated colder zone where fuel temperature may decrease. This explains why fuel decomposition and coke formation are observed despite the measured heated fuel temperature is always lower than 850 K (which is the minimum required one to get ethylene decomposition [45,46]). Hence, Eq. (10) may lead to an underestimation of the real sensible heat flux absorbed by the fuel heated by the flame. The chemical heat flux qfchem , cc is calculated as follows:

qchem f,cc = mf

Hpyr

pyr

3. Results and discussion 3.1. Influence of fuel mass flow rate and equivalence ratio on cooling efficiency An accurate characterization of the flame has been provided in previous works, where a detailed analysis of the temperature field characterizing the combustion gases has been presented and the coupling between fuel decomposition and fuel combustion deeply investigated [35,44,48]. In this section, we detail the effects of fuel mass flow rate ṁf and equivalence ratio φ variations on: i) the sensible and the chemical heat fluxes absorbed by the fuel in the cooling channel, ii) the heat flux density passing from the combustion gases to the combustor wall, and iii) the heat transfer efficiency of the cooling system. To investigate the influence of ṁf and φ on these heat fluxes, test cases 2, 3 and 4 are presented (Table 1). These tests have been carried out by varying ṁf between 2.0 and 2.8·10−5 kg.s−1, with φ respectively equal to 1.0, 1.25 and 1.5. They have already been partially examined in [35], where the effect of ṁf and φ variations on the temperatures of the combustion gases and on that of the decomposed fuel have been detailed. In Figs. 3 and 4 the sensible and the chemical1 heat fluxes absorbed by the fuel at steady state, divided by the internal surface of the combustion chamber, are given for these three experiments for steady state from 1 to 5. As explained in [35], for test 4 (φ = 1.5) data are only available for steady state plateaus from 1 to 3, because during this test the decomposition of the fuel originated coke deposits that have caused the occlusion of the cooling channel and the interruption of the test before its end. As mentioned in Section 2.3, the sensible heat flux is calculated using Eq. (10), even if probably temperature Thf, is not the highest one achieved by the coolant in the channel. This can be seen in Fig. 5, presenting the variation of the temperature of the fuel along the cooling channel, obtained with the CHEMKIN-PRO calculations, for steady state 2. This figure indicates that, depending on the case, the maximum temperature achieved by the coolant can be even 200 K higher of the one measures at the cooling channel outlet. Hence, the use of Eq. (10) leads to underestimate the real absorbed sensible heat load. Yet, it represents a first method to evaluate the performance of the cooling system and to compare tests run under different conditions. The increase of qfsens , cc with fuel mass flow rate and equivalence ratio cannot be justified by a variation of fuel residence time tf in the cooling channel. Indeed, as it can be observed in Table 2, presenting the variation (obtained with the CHEMCHIN-PRO simulations) of tf as a function of ṁf for test case 2, 3 and 4, in general fuel residence time decreases with ṁf and φ. The rise in this heat flux with fuel mass flow rate is in part because, as shown for test case 3 in Fig. 6, an increase in ṁf leads to an increase in the Reynolds number Ref of the fuel along the cooling channel. As a growth of Ref leads to a higher heat transfer rate between the channel internal surface and the fuel, qfsens , cc increases occurring after ṁf increases are at least partially due to the consequent augmentation of heat transfer rate between the channel wall and the fuel. Nevertheless, it is not easy to define the appropriate weight to accord to this parameter among others. In fact, as it can be observed in Fig. 7 for tests 2, 3, and 4, an increase in equivalence ratio, leading to an increase in qfsens , cc , is also accompanied by a drop in Ref along the cooling channel. The augmentation of qfchem , cc with fuel mass flow rate and equivalence ratio is due to the increase, with both ṁf and φ, in the decomposition rate of the fuel. This can be seen in Table 3, giving ethylene pyrolysis rates for test cases 2, 3, and 4 for steady states from 1 to 5. Table 3

(11)

where τpyr and ΔHpyr are respectively fuel pyrolysis rate and fuel decomposition enthalpy, both obtained with the CHEMKIN-PRO numerical calculations mentioned above in this section. To characterize the performance of the combustor in terms of heat recovery, the heat transfer efficiency η of the combustor seen as a counter-flow heat exchanger has been calculated. It is defined as the ratio between the actual rate of heat transfer from the combustion gases to the coolant and the optimum one [47]:

=

qf,cc qf,cc |opt

=

qf,cc U Aiw

(12)

Tlog

where:

U=

Dw,int ln(Dw,ext /Dw,int ) 1 1 1 + + + hg hr 2 k cc hf

Tlog =

(TTC2 g log((TTC2 g

Thf ) Thf )

(TTC11 g (TTC11 g

1

(13)

Tcf ) Tcf ))

(14)

In Eqs. (12)–(14), hg and hr are respectively the convective and the radiative heat transfer coefficient between the burned gases and the combustor internal surface, Aiw is the combustor internal surface, hf is the convective heat transfer coefficient between the cooling channel internal surface and the fuel, kcc is the cooling channel thermal conductivity, TgTC2 and TgTC11 are the average gas temperature between positions P1, P2, and P3 measured respectively at combustor basis and at combustor outlet.

1

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Fig. 3. Sensible heat flux density absorbed by the fuel at steady state for tests 2, 3, and 4.

Fig. 6. Fuel Reynolds number along the cooling channel at steady state 1, 2 and 3 for test 2.

Fig. 4. Chemical heat flux density absorbed by the fuel at steady state for tests 2, 3, and 4.

Fig. 7. Fuel Reynolds number along the cooling channel at steady state 2 for tests 2, 3, and 4. Table 3 Fuel pyrolysis rate at steady state for test cases 2, 3, and 4.

Table 2 Fuel residence time in the cooling channel at steady state for test cases 2, 3, and 4. φ

tf|1 (s)

tf|2 (s)

tf|3 (s)

tf|4 (s)

tf|5 (s)

2 3 4

1.0 1.25 1.5

4,78 1,91 1,98

4,50 1,88 1,86

4,45 1,71 1,59

4,88 1,89 –

4,73 2,04 –

φ

2 3 4

1,0 1,25 1,5

p |1

0,01% 0,65% 3,72%

p |2

0,03% 0,72% 17,6%

p |3

0,24% 2,06% 36,1%

p |4

0,05% 1,43% –

p |5

0,01% 0,48% –

higher for test case 4 than for the others, as it can be observed in Table 4, which presents for these tests the ratio between qfchem , cc and qf , cc at steady state. The rise in pyr as a function of ṁf and φ is mainly because an increase in these parameters causes an increase in the temperatures achieved by the fluid in the cooling channel (see Fig. 5 above for the effect of equivalence ratio and Fig. 8. for that of fuel mass flow rate). This enhances the decomposition of the fuel, which is highly dependent on temperature [7]. To identify the parameter that influences the most the sensible and the chemical heat fluxes absorbed by the fuel, a sensitivity analysis has been carried out. Results are given in Figs. 9 and 10, indicating that, chem between ṁf and φ, the parameter affecting qfsens , cc and qf , cc the most is equivalence ratio. To conclude the analysis of the heat flux absorbed by the decomposing fuel, it is worth noting that a hysteresis effect can be observed both in Figs. 3 and 4, by comparing respectively steady states 1 and 5 and steady states 2 and 4. This confirms what was said in [35]. In Figs. 11, 12, and 13, the total, the convective, and the radiative heat flux densities qg, qg,conv and qg,rad passing from the combustion gases to the combustor wall at steady state for experiments 2, 3, and 4 are given. Figs. 11, 12, and 13 indicate that in general, independently

Fig. 5. Fuel temperature along the cooling channel at steady state 2 for tests 2, 3, and 4.

Test case

Test

indicates that ethylene pyrolysis rate is, for test case 4, about one order of magnitude higher than for test case 3 and about two orders of magnitude higher than for test case 2. Consequently, the chemical heat sink of the fuel-coolant is far 1096

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Table 4 Ratio between qfchem , cc and qf , cc at steady state for test cases 2, 3, and 4. Test

φ

qchem f, cc / q f, cc |1

qchem f, cc / q f, cc |2

qchem f, cc / q f, cc |3

qchem f, cc / q f, cc |5

qchem f, cc / q f, cc |5

2 3 4

1,0 1,25 1,5

1,64% 2,05% 9,78%

0,12% 2,06% 27,2%

0,38% 5,21% 39,5%

0,21% 4,01% –

0,01% 1,43% –

Fig. 11. Total heat flux density qg, from the combustion gases to the combustor wall at steady state for tests 2, 3, and 4.

Fig. 8. Fuel temperature along the cooling channel at steady state from 1, 2, and 3 for test 4.

Fig. 12. Convective heat flux density qg,conv, from the combustion gases to the combustor wall at steady state for tests 2, 3, and 4.

Fig. 9. Percentage change of qfsens , cc as a function of ṁf and φ for tests 2, 3, and 4.

Fig. 13. Radiative heat flux density qg,rad, from the combustion gases to the combustor wall at steady state for tests 2, 3, and 4.

Fig. 10. Percentage change of qfchem , cc as a function of ṁf and φ for tests 2, 3, and 4.

and the combustor wall. Consequently, both convective and radiative heat transfer from the burned gases raise. According to Fig. 11, the total heat flux density qg increases when equivalence ratio φ passes from 1.0 to 1.25 whereas a further raise of φ from 1.25 to 1.5 seems to produce a drop. This appears to be

from equivalence ratio, these three fluxes increase with ṁf. This is because, as shown for test 2 in [35], an increase in fuel mass flow rate causes a rise in the temperature gradient between the combustor axis 1097

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inconsistent with 3 and 4, which indicate that an increase in equivalence ratio from 1.25 to 1.5 causes a strong increase in the heat flux absorbed by of the coolant. As observed, this increase cannot be justified by tf and Ref variations because both parameters decrease when equivalence ratio raises from 1.25 to 1.5. Hence, it must be the result of an increase in the heat lux generated by the burned gases.

Table 5 Calculated flame emissivities at steady state for test cases 2, 3, and 4. Test

φ

εg

2 3 4

10 1.25 1.25

0.065 ± 0.004 0.123 ± 0.018 0.263 ± 0.040

Table 6 Combustor heat transfer efficiencies at steady state for test cases 2, 3, and 4. Test

φ

η|1

η|2

η|3

η|4

η|5

2 3 4

1,0 1,25 1,5

0,081 ± 0,013 0,096 ± 0,014 0,084 ± 0,014

0,095 ± 0,014 0,109 ± 0,016 0,095 ± 0,018

0,107 ± 0,016 0,118 ± 0,018 0,107 ± 0,021

0,089 ± 0,013 0,105 ± 0,016 –

0,077 ± 0,012 0,091 ± 0,014 –

Fig. 14. Time-averaged carbon coking rate, time-averaged fuel temperature in the cooling channel and time-averaged heat flux density generated by the combustion gases as a function of the distance from the inlet of the channel and from the bottom of the combustor for test 1.

Fig. 15. Time-averaged carbon coking rate, time-averaged fuel temperature in the cooling channel and time-averaged heat flux density generated by the combustion gases as a function of the distance from the inlet of the channel and from the bottom of the combustor for test 8.

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Fig. 16. Fuel pressure drop in the cooling channel as a function of time during test 8.

The drop of qg when φ passes from 1.25 to 1.5 can probably be explained by the fact that, under fuel-rich conditions, fuel combustion forms soot deposits that accumulate on the thermocouples measuring the temperature of the combustion gases when, at a steady state, they are placed on the axis of the combustor [49,50]. Thus, due to the isolating power of soot, the temperatures measured on this axis are underestimated and, consequently, the calculated heat transfers qg,conv and qg,rad are underestimated too. This underestimation is greater when φ = 1.5 than when φ = 1.25 because, as shown in [35], under this condition soot formation is enhanced. The formation of soot particles also explains why heat transfer from the burned gases is enhanced under fuel rich conditions. In fact, the presence of carbon particles increases flame luminosity and emissivity. Hence, radiative heat transfer between the burned gases and the combustor wall, strongly depending upon flame emissivity, raises [39]. This can be clearly observed in Fig. 13, indicating that qg,rad strongly increases as a function of φ. The associated average flame emissivities εg are given in Table 5. Fig. 12 shows a decrease in convective heat transfer qg,conv with equivalence ratio. This is probably because, differently from the radiative one, the convective heat transfer between the combustion gases and the combustor wall is directly influenced by burned gases speed, which decreases with φ. It is also worth noting that qg variations with ṁf that can be observed in Fig. 11 appears to be are higher under stoichiometric combustion. Therefore, from an engineering perspective, when designing the cooling system, increasing the thrust under rich conditions should impose a lower additional heat flux evacuation by cooling than under stoichiometric ones. To define the performance of the regeneratively cooled combustor, the heat transfer efficiency η of the overall cooling system has been determined for test cases 2, 3, and 4 (Table 6). It can be observed that η grows as a function of ṁf whereas, with respect to φ, a maximum is reached at 1.25. A hysteresis effect can be noted by comparing test 2 and test 3. Indeed, it can be seen that for both experiments the values reached by η at steady state 4 and 5 are always lower than those reached respectively at steady state 2 and 1. It is worth noting that being the sensible (and consequently the total) heat flux absorbed by the pyrolyzing fuel underestimated (as explained above), the calculated heat transfer efficiencies are underestimated too.

3.2. Coking activity monitoring methods suitable for real-time on-board application As already observed, fuel thermal stability is a great challenge facing the use of endothermic hydrocarbon fuels. To investigate the consequences of fuel coking activity on the cooling system, two test cases (1 and 8) are examined. Both tests were stopped before the end, because the coke deposits generated by fuel decomposition had caused the occlusion of the cooling channel. Figs. 14 and 15 present, respectively for test 1 and 8, the timeaveraged carbon deposition rate in the cooling channel, the timeaveraged fuel temperature and the time-averaged heat flux density generated by the combustion gases as a function of the distance from the inlet of the channel and from the bottom of the combustor. To calculate the time-averaged carbon formation rate, at test completion the rolled-up tube was removed from the combustor, washed with hexane and dichloromethane and dried with nitrogen. It was then cut into 3 m long pieces and the carbon accumulated in each section weighted with a precision balance (Kern ABS-N/ABJ-NM Analytical Balance). To express fuel coking rate in kg.m−2.s−1, the measured coke mass has been divided by cooling channel internal surface and by test duration. This approach could be questionable for two reasons. First, carbon coking activity is not uniform all over the experiment, as the temperatures achieved by the decomposing fuel vary with time. Second, it probably leads to underestimate the real carbon deposition rate, as during the first part of the experiment the temperatures achieved by ethylene in the cooling channel are too low to allow its pyrolysis. Nevertheless, it permits to obtain consistent results and to compare experiments run under different conditions. Figs. 14 and 15 indicate that carbon deposits only form in the highest temperature sections of the cooling channel, where the temperature achieved by the hydrocarbon is over 800 K, and that fuel deposition reaches its maximum at the bottom of the combustion chamber, where the temperature and the heat flux are the higher. This outcome is consistent with the mainstream literature works, which states that, when few seconds residence times are considered, hydrocarbon pyrolytic deposition may only occur at temperatures over 800 K [7,31,33]. The extent to which regenerative cooling benefits can be capitalized is not only related to the ability to moderate coke formation, but also to the development of carbon deposition monitoring methods suitable for real-time on-board application. In this sense, the monitoring of fuel1099

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coolant pressure drop in the cooling channels (ΔP) could be a viable strategy. Fig. 16 gives ΔP as a function of time for test 8. This parameter is calculated as the difference between the pressures measured by pressure transducers PTE-1 and PTE-2 respectively (see Fig. 1). The two very abrupt pressure loss increases indicated as ΔP1 and ΔP2 result from fuel mass flow rate raises (respectively from 2.0 to 2.4·10−5 kg.s−1 and from 2.4 to 2.8·10−5 kg.s−1), whereas the brusque pressure loss decrease indicated as ΔP3 is due to fuel flow rate reduction from 2.4 to 2.0·10−5 kg.s−1. Instead, the progressive pressure drop increases ΔPb1 and ΔPb2 are both due to the gradual accumulation of coke deposits on the cooling channel internal surface, which reduces the tube cross section. As already mentioned, this test was run until the final sudden fuel pressure drop raise that can be easily seen in the figure, indicating that complete blockage was imminent.

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4. Conclusions Regenerative cooling is expected to permit the thermal protection of Scramjet propelled vehicles flying at velocities over Mach 5. Its effectiveness is studied in this paper when an endothermic hydrocarbon fuel is used. A regeneratively cooled combustor, allowing the experimental study of pyrolysis-combustion coupling and fuel coking activity, has been designed. It is used to run tests under both stationary and transient conditions using ethylene as fuel and air as oxidizer. The influence fuel mass flow rate and equivalence ratio, on the heat flux between the burned gases and the combustor internal surface, on the heat flux absorbed by the fuel in the cooling channel and on combustor heat exchange efficiency have been determined. Under rich conditions, radiative heat transfers from the combustion gases to the combustor wall are higher than the convective ones, being the situation the opposite at stoichiometric conditions. Both the sensible and the chemical heat fluxes absorbed by the fuel in the cooling channel rise with fuel flow rate and equivalence ratio. The sensibility of these heat fluxes to these two parameters has been determined. It has also been observed that combustor heat exchange efficiency increases with fuel mass flow rate whereas, with respect to equivalence ratio, it has a maximum. Fuel coking activity has been studied. A fuel coking activity monitoring method suitable for real-time on-board application, i.e. the measure of fuel pressure drop in the cooling channels, has been tested and validated. The knowledge acquired in this paper represents a first step to better understand the effect of the command parameters on the main outputs characterizing a regeneratively cooled combustor. It will now be possible in a future work to develop analytical relationships for an engine control strategy to generate a required thrust without exceeding a temperature threshold. References [1] Curran ET. Scramjet engines: the first forty years. J Propuls Power 2001;17:1138–48. https://doi.org/10.2514/2.5875. [2] Van Wie DM, D’Alessio SM, White M. Hypersonic air breathing propulsion. Johns Hopkins APL Technical Digest 2005;26:430–6. https://doi.org/10.2514/4.470356. [3] Van Griethuysen VJ, Beach WP, Petleyt DH. High-speed flight thermal management. Developments in high-speed vehicle propulsion systems. Progr Aeronaut Astronaut 1996;165:517–79. https://doi.org/10.2514/5.9781600866401.0517.0579. [4] Daniau E, Sicard M. Experimental and numerical investigations of an endothermic fuel cooling capacity for scramjet application. In: 13th AIAA/CIRA international space planes and hypersonics systems and technologies. Capua. May 16–20 2005: doi:10.2514/6.2005-3404. [5] Ulas A, Boysan E. Numerical analysis of regenerative cooling in liquid propellant rocket engines. Aerosp Sci Technol 2013;24:187–97. https://doi.org/10.1016/j.ast. 2011.11.006. [6] Hou LY, Dong N, Sun DP. Heat transfer and thermal cracking behavior of hydrocarbon fuel. Fuel 2013;103:1132–7. https://doi.org/10.1016/j.fuel.2012.09.021. [7] Maurice L, Edwards T. Liquid hydrocarbon fuels for hypersonic propulsion. Scramjet propulsion. Progr Aeronaut Astronaut 2001;189:757–822. https://doi. org/10.2514/5.9781600866609.0757.0822. [8] Heiser HH, Pratt DT. Hypersonic air-breathing propulsion 1994;26. doi:10.2514/4.

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