Autoignition of a liquid n-heptane jet injected into a confined turbulent hot co-flow

Hannah Moran, Ajay Gupta, Victor Voulgaropoulos, Christos N. Markides*

Clean Energy Processes (CEP) Laboratory, Department of Chemical EngineeringImperial College London, London, United Kingdom

email: c.markides@imperial.ac.uk

ABSTRACT

Alternatives to conventional combustion engines, such as gasoline direct injection engines, homogeneous charge compression injection engines and dual-fuel turbines, promise improved fuel efficiency and reduced emissions. The present study of liquid-fuel autoignition in turbulent flows explores the underlying phenomena in these applications towards next-generation combustors. Experiments have been performed on the autoignition of continuous liquid n-heptane jets injected axisymmetrically into confined turbulent coflows of preheated air. Jet atomisation was characterised using high-speed imaging, and autoignition locations and corresponding delay times were recorded for various bulk air temperatures and velocities. Two turbulence-generating plates with different perforation sizes were used to investigate the role of turbulence in affecting the phenomena under investigation. Smaller droplets formed in flows with lower turbulence intensities and larger integral lengthscales. The autoignition length increased and delay time decreased with increasing bulk air velocity, the latter being contrary to results from pre-vaporized n-heptane autoignition in an identical apparatus.

KEYWORDS

Autoignition; Turbulence; Two-phase combustion; Liquid fuel; Delay time; Reacting flow

INTRODUCTION

The autoignition of liquid fuels in turbulent air flows has wide-ranging applications in the fields of clean combustion and efficient fuel consumption. The improvement of diesel and gasoline direct injection (GDI) engines enables more efficient combustion, reducing fossil fuel use and related emissions. Homogeneous charge compression injection (HCCI) engines offer reduced NOX and soot emissions, and provide a possible alternative to conventional diesel combustion [1]. Dual fuel turbines facilitate fuel flexibility and provide an efficient and reliable power supply for stationary power generation applications, and the use of dual-fuel mode in diesel engines decreases NOX, CO2 and particulate matter emissions [2].

Investigations of the autoignition of gaseous fuels in hot turbulent air flows found that the autoignition delay times and autoignition lengths increased with bulk air velocity [3], but decreased with a simultaneous decrease in turbulence intensity and increase in turbulent lengthscale [4]. While it is expected that turbulence would promote mixing and hence accelerate autoignition, these recent experimental findings and also complementary DNS [5] suggest that it can instead have a delaying effect on the phenomenon. Therefore, the chemical and mixing processes cannot be decoupled in the treatment of inhomogeneous autoignition.

* Corresponding author

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When a liquid, rather than a gas, fuel is used, the autoignition process becomes a multiphase problem which is significantly more complicated than the gaseous fuel autoignition. Two-phase autoignition involves jet break-up, atomisation and evaporation of the liquid fuel, and further mixing and reaction of the resulting vapour with the surrounding air. Droplet and spray vaporisation and combustion have been the subject of numerous experimental and numerical studies, which have been recently been reviewed [6-10]. In contrast to the gas fuel autoignition behaviour, Mizutani et al. [11] found that turbulent mixing shortened the autoignition delay time and concluded that the autoignition process was controlled by the droplet evaporation rate and not the chemical reaction kinetics. Hinkeldey et al. [12] undertook an experimental investigation of fuel droplet-laden hot air flow at high pressures and observed that autoignition occurred at random locations within the reacting flow, with autoignition spots resulting in flames that expanded with downstream convection and consumed residual droplets in the flow.

Kourmatzis & Masri [13] investigated the effects of grid generated velocity fluctuations on the primary atomisation and subsequent droplet deformation of non-reacting sprays with a cold air coflow, namely laminar liquid methyl esters, acetone and diesel jets. Velocity fluctuations were found to randomise the break-up process. Further, O’Loughlin & Masri [14,15] studied the autoignition characteristics of dilute turbulent spray flames of ethanol and methanol formed in hot coflowing stream of vitiated combustion products. Similarly, Rodrigues et al. [16] performed experiments with ethanol pressure-swirl spray flames in a coflow of either air or hot-diluted oxidant. They observed that the liquid jet was immediately disrupted on introduction to the hot-diluted coflow. The atomisation mechanism in hot-diluted coflow was different to that in pure ambient cold air coflow. Gordon & Mastorakos [17] investigated the effects of flow on the autoignition of dilute diesel and biodiesel sprays injected at right angles to a hot air turbulent flow as monodisperse individual droplets. They found that the time-averaged autoignition length increased with increasing air velocity and turbulence intensity.

The objective of this work is to understand the effect of turbulence on the atomisation, evaporisation, mixing and autoignition of liquid fuel jets in turbulent hot air flows. The effect of turbulence is investigated experimentally by considering the effect of different air velocities and grid perforation sizes in the turbulence generating plate on the droplet size, autoignition length and autoignition delay time. Results for each pair of variables are reported, and provide an insight into the autoignition phenomena. These results constitute a valuable data set for development and validation of computation autoignition models.

EXPERIMENTAL METHODS

The apparatus used for the experiments in this study was identical in terms of geometry, but with some further developments to allow liquid fuel injection, autoignition and combustion, to that employed by Markides and Mastorakos [3] for the investigation of gaseous fuels in turbulent flows, referred to as the ‘confined turbulent hot co-flows (CTHC)’ apparatus and represented schematically in Figure 1. A detailed description of the apparatus is provided by Gupta [18], whilst information pertinent to these experiments is given below.

Hot air preparationA Bronkhorst digital mass-flow controller (air flow range of 0.5-24.1 kg/h) was used to control and meter the flow of dried and filtered compressed air, which was heated by an Orsam Sylvania industrial-grade 8 kW high temperature inline heater (F076619) to temperatures up to 1150 K. A PID controller was used to control the power input to the heater through a thyristor assembly, with a feedback input of a measurement from a welded tip

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twisted pair K-type thermocouple at the air heater exit. On leaving the electric heater the hot air first passed through a stainless steel perforated plate with 1 mm diameter holes and 75% solidity into a stainless steel hot air adapter. Downstream of this, the air passes through a turbulence generating plate of hole diameter M and a 60° hole configuration. Two sizes of the perforated plate were used in these experiments:

1. Large perforated plate (LPP): M = 3.0 mm, solidity = 42%, pitch = 3.7 mm,2. Small perforated plate (SPP): M = 1.2 mm, solidity = 55%, pitch = 1.7 mm,

where pitch is defined as the centre-to-centre spacing of adjacent holes. If zis defined as the axial direction coordinate, where z = 0 at the fuel injection location, the perforated place was located at z=−¿63 mm. A 10 cm thick layer of refractory high-temperature insulation wool was used to heavily insulate the hot-air adapter and heater. In order to monitor the thermal state of the CTHC apparatus during the heating up process a 1.5 mm diameter mineral-insulated N-type thermocouple was embedded into the hot-air adapter wall to confirm that thermal equilibrium was obtained before performing the experiments.

Quartz test sectionFrom the adapter, the hot air entered an optically accessible and vertically mounted custom-made quartz test-section consisting of a 65 cm long central quartz tube, subsequently referred to the ‘quartz reactor’ and abbreviated to ‘qzr’. The reactor had an inside diameter dqzr=¿25 ± 0.2 mm and an outside diameter Dqzr=¿28 ± 0.2 mm and operated at atmospheric pressure since it was open-ended at the exit, exhausting into the laboratory exhaust fume. It was enclosed along its entire length by a quartz jacket of square cross section, the shape of which was selected to minimise optical distortion due to curved surfaces. To reduce heat losses from the quartz reactor to the ambient, a counter flowing double-pipe type heat exchanger was set up, where hot air flowed from a second process air heater horizontally mounted hear to the exit of the reactor and exited through a port near to the base of the test section.

Liquid fuels and injection systemThe liquid fuel used was pure n-heptane (99+% pure Acros Organics from Fisher Scientific, CAS number: 142-82-5). A water-cooled liquid-fuel injector nozzle was developed in house

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Figure 1. Schematic of the CTHC apparatus. Figure is not to scale.

and comprised three concentrically arranged stainless steel hypodermic tubes 11G (ID: 2.388 ± 0.076 mm, OD: 3.048 ± 0.025 mm), 17G (ID: 1.067 ± 0.038 mm, OD: 1.473 ± 0.013 mm), and 24G (ID: 0.311 ± 0.019 mm, OD: 0.565 ± 0.006 mm); G stands for gauge. The liquid fuel was carried in the innermost 24G tube, and coflowing cooling water from the laboratory mains flowed in the parallel 17G tube and was returned to the laboratory environment by the 11G tube. The injector’s surface was coated by insulating high-temperature alumina cement at all points exposed to hot air flow. The outside wall temperature of the 24G tube was measured by a 0.25 mm diameter mineral insulated N-type thermocouple. This measurement was then used as an estimate of the fuel temperature at injection, T fuel, which was influenced by the temperature of the hot coflowing air and was therefore not an independent process variable. A 100 mL stainless steel syringe mounted on a programmable Harvard Apparatus PHD Ultra syringe pump was used to deliver the liquid fuel to the injector nozzle, with both items mounted vertically to facilitate bleeding of any vapours into the fuel line. The syringe was refuelled from a 250 mL borosilicate separating funnel functioning as a fuel reservoir.

Key experimental variablesThree key measurements are reported for each autoignition experiment, as described below:

1. The bulk air inlet temperature,T air, was measured at the quartz reactor inlet (z=−¿40 mm) with a 0.25 mm bare wire butt-welded R-type thermocouple (Pt/13%Rh-Pt). A custom-made probe assembly was used to embed this thermocouple into the wall of the hot air adapter such that its butt-welded junction was located at 9 ± 1 mm into the flow, i.e. r /r qzr= 0.3, where r is the radial location. All thermocouple measurements were recorded with a 6-bit multifunction National Instruments USB-6259 data acquisition system with SCB-68 connector blocks. The raw temperature measurements of the R-type thermocouple were corrected for conduction and radiation losses as per Appendix A of Markides [19] to obtain T air. It was found that for T air between 1060 K and 1135 K, (i) at Uair between 17.5 and 25 m/s, the total correction was ± 4 K i.e., 0.4% of T air; (ii) at Uair between 25 and 37 m/s total correction was ± 3 K i.e., 0.3% of T air.

2. The bulk air inlet velocity, Uair, was based on the mass flow rate of air, mair, as measured by the air mass flow controller and was defined as the average velocity over the annular cross section area between the fuel injector nozzle and the air mass flow controller. Therefore, for a given T air,

Uair=mair

ρair [( π4

dqzr2 )−( π

4Dinj

2 )] (1)

3. The liquid fuel injection velocity, U fuel, was defined as the average velocity over the cross-sectional area of the 24G hypodermic tube based on the volumetric flowrate measured directly at the syringe pump and is given by:

U fuel=Qfuel ρfuel

amb

ρ fuel [ π4

d inj2 ] (2)

The thermophysical properties of n-heptane were obtained from NIST [20].4. The turbulence-generating plate hole size, M , was also varied, as described previously.

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Turbulence characterisationOf interest in turbulent flows are the local velocity field and in particular the velocity fluctuations relative to a mean velocity, where the fluctuations can be characterised by the normalised root mean square (RMS) of the axial velocity fluctuations, u '. The turbulent intensity is defined as the ratio of the velocity fluctuations to the local mean axial velocity, i.e. u' / ⟨ uz (r )⟩. A longitudinal integral lengthscale, Lturb, can also be obtained by integrating over the normalised autocorrelation function of the axial velocity fluctuations and using Taylor’s hypothesis. Figure 2 shows the radial profiles of the mean and RMS of the local axial velocities under non-reacting conditions for the CTHC apparatus.

The CTHC apparatus configuration presented in this work was identical to that used for previous investigation of autoignition of gaseous fuels [3,4,19]. For these studies, a hot wire and a constant temperature anemometer (CTA) system were used to characterise the air velocity field. Similar air velocity characteristics were expected in the quartz reactor of this work, particularly at the inlet and until this was affected by the liquid fuel evaporation and reaction. It was reported that with the LPP, the axial velocity profile was uniform over r /r qzr≅ 0.15 – 0.8 for z= 2 – 42 mm; the normalised mean axial velocity ⟨ uz (r ) ⟩ /U air = 1.10 – 1.14. A 2-mm thick velocity boundary layer grew along the walls of the quartz reactor and the fuel injector nozzle. The flow turbulence intensity, u' / ⟨ uz (r )⟩ was 0.11 – 0.14 for 0.4 < r /r qzr < 0.7 at z = 2 mm while Lturb was 3.5 ± 1 mm. For SPP, Lturb was 5 – 7 (± 1) mm at z = −10 mm [4]. In both cases, Lturb was of the same order as M .

Further work on the CTHC apparatus [21,22] found that at any given spatial location increasing Uair resulted in an increase in the velocity fluctuations, u ', and a decrease in Lturb .

Thus, different values of U air and M lead to flows with different turbulence characteristics, and the effect of turbulence on autoignition can be studied by varying Uair and M (by using either the LPP or SPP) as done in earlier work in a similar apparatus [3].

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Figure 2. Radial profiles of the mean and RMS of the local axial velocities under non-reacting conditions at ambient (293 K) temperature. The mean local axial velocity normalised relative

to the bulk air velocity, ⟨uz (r ) ⟩ /U air, is presented on the left. The turbulent intensity, i.e. normalised RMS velocity fluctuations relative to the local mean axial velocity u' / ⟨ uz (r )⟩, is

presented on the right for the flow condition ℜ = 6160.

Autoignition characterisationTo study the effect of turbulence on autoignition, the definition of quantifiable parameters that characterise autoignition is required; here, we use the autoignition length and delay time. In this work, the earliest or most upstream location of autoignition was considered, in the random spots regime only, as this was a true representation of the autoignition phenomenon and was not affected by heat released by post-ignition propagating flames. To quantify this, the 5th percentile location L5 pc was selected from the axial distribution of the autoignition locations. This served as a statistically robust and good estimate of the furthest upstream autoignition location.

An autoignition delay time can be defined as the mean residence time, τ5 pc, from injection to autoignition based on the L5 pc such that τ5 pc=L5 pc /U air, where Uair is the characteristic velocity of the process. The delay period represented by τ5 pc consists of both a physical delay during which liquid-fuel atomisation, evaporation and mixing with air occur, and a chemical delay during which pre-ignition reactions develop [23].

OperationHeating up of the apparatus was first performed for approximately 90 minutes to reach the desired operation temperature, T air, while thermal equilibrium was attained by running the heaters for approximately another 45 minutes. The flow rates of the air and the fuel were then set and measurements were taken after a few minutes to ensure steady-state conditions.

Optical measurements & image processingThe primary atomisation and autoignition phenomena of the liquid fuel were captured with high-speed imaging. An iSpeed 3 (Olympus) camera equipped with a Nikkor 85mm lens was used at full resolution of 1280x1024 pixels2 and at an acquisition frequency of 2 kHz. For the primary atomisation phenomena, two extension rings amounting to 32 mm, along with a LED floodlight combined with a polystyrene light diffuser were employed, while a short exposure time of 50 μs was utilised to ensure image sharpness. The autoignition phenomena were captured by chemoluminescence imaging and hence a higher exposure time of 250 μs was set, in order to ensure that enough light passes to the camera sensor. 104 images were obtained for each experimental run to establish statistical convergence for the measurements.

Image calibration was performed with a graticule target placed inside the reactor [24]. For the atomisation analysis, segmentation was performed with a rolling pixel-wise temporal median filter and the GrabCut algorithm [25] that is based on the pixel intensity gradients between the droplets and the medium. A similar post-processing approach was also used by Voulgaropoulos and Angeli [26] for droplet tracking and sizing. Additionally, a fully-convolutional neural network [27] was trained to predict the locations of the droplet centres, and to separate out the droplets in contact with each other. For the autoignition measurements, the background was separated with a rolling temporal median filter, and the image was then binarised and the continuous regions were characterised as reaction zones. Temporally-resolved tracking of each reaction zone was conducted with the Kanade-Lucas-Tomasi algorithm, already implemented in MATLAB. For more details on the optical measurements and the image processing steps for both the atomisation and the autoignition analysis, the reader can refer to Gupta [18].

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RESULTS AND DISCUSSION

A series of experimental runs were performed in which a constant T air was maintained at various values in the range of 1107 – 1138 K, U fuel was kept constant at either 0.7 m/s or 1.1 m/s, and U air was increased in the range of 14 – 34 m/s. The effect of each perforated plate on the structure of turbulence was investigated.

Droplet sizeUnder all experimental conditions primary atomisation of the liquid fuel was complete before z = 20 mm, indicating that this was the most upstream location at which liquid-fuel drop existed. The effect of Uair and M on the mean droplet diameter, dmean, was investigated. It was found that dmean increased monotonically with increasing z due to droplet coalescence. Values for dmean at z > 22 mm are excluded in the results, as a parallel analysis suggested that the overall effect of reactor conditions on dmean downstream of this remained the same.

Effect of air velocity. Figure 3 shows the effect of Uair on dmean at various T air and U fuel conditions with both the LPP and SPP. In general, dmean decreased with increasing U air, although this effect was more pronounced for the LPP than the SPP. For the LPP, over an increase in U air from 15 m/s to 20 m/s, dmean decreased from approximately 800 m to 600 m at U fuel = 0.7 m/s, whilst at U fuel = 1.1 m/s, dmean decreased from approximately 650 m to 600 m. The comparable changes for the SPP are considerably small, with sizes of approximately 50 m and 10 m respectively.

Effect of grid-hole size. The effect of the turbulence generating grid perforation size, M , on dmean can be observed in Figure 3. At both values of U fuel investigated, the droplets were smaller with the SPP (smaller M ) than the LPP for similar T air and U air conditions. For the condition U fuel = 1.1. m/s, the dmean values for the droplets with the SPP were 30% smaller than with the LPP, and the standard deviations were 40% smaller. The larger Lturb and lower uair

' with the SPP can result in liquid jet thinning, and these observations of dmean suggest that these longer and thinner liquid jets atomise into smaller droplets. This effect has been

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(a) (b)

Figure 3. Effect of U air on dmean at U fuel = 0.7 m/s (a) and 1.1 m/s (b) at six sets of constant T air conditions for two grid sizes, LPP M = 3 mm in empty markers and dash-dot lines, SPP M =

1.2 mm in filled markers and solid lines.

observed in other studies: Eggers & Villermaux [28] reported that thinner, stretched and corrugated liquid ligaments produced droplets of a smaller size and a narrower size distribution; Kourmatzis & Masri [29] observed the formation of smaller droplets with increasing jet thinning for a finer grid.

Autoignition lengthEffect of air velocity. The effect of air velocity on autoignition length for the LPP and SPP is presented in Figure 4. Over the presented conditions L5 pc increased with increasing Uair, i.e. the autoignition region moved downstream. With the LPP, for U air in the range 14.4 m/s to 27.5 m/s and T air up to 1133 K, the downstream shift was 6 – 16 mm per m/s of Uair, whilst decreasing Uair over 27.5 m/s resulted in a decrease in L5 pc . Similar behaviour was observed with the SPP, where L5 pc also shifted upstream at Uair > 27.5 m/s for U fuel = 1.1 m/s.

In the previous section, it was shown that dmean decreased with increasing U air. This results in an increased droplet surface area, and therefore increased heat transfer area, causing evaporative cooling of the air flow within the quartz reactor. This goes some way to explaining the downstream shift of the autoignition region up to a certain Uair, but the upstream shift of L5 pc at higher U air cannot be understood in this way.

For the autoignition of gaseous n-heptane in a heated turbulent air flow that was established in a similar apparatus [3], autoignition was observed to shift non-linearly downstream with an increase in U air at constant T air [30]. This retardation of autoignition cannot, however, be explained by evaporative cooling, suggesting that the increased turbulent mixing is inhibiting autoignition. This behaviour was also observed by Markides [19].

Effect of perforation size. Experimental results for the autoignition length with the LPP and SPP are presented in Figure 4. The effect of T air and Uair on L5 pc was qualitatively the same for both the LPP and SPP, and both sets of results exhibited similar rates of decrease in L5 pc with Uair . However, in general L5 pc with the SPP was 20 – 50 mm shorter than with the LPP, indicating that autoignition occurred closer to the injector nozzle.

It has already been established that dmean decreased with the SPP. By the argument of increased evaporative cooling due to increased surface area the L5 pc would be expected to be longer with the SPP. However, this was not the case in the current experiments, which highlights the effect of the different flow turbulence characteristics of the two plates. Also, if it is considered that the SPP was characterised by a larger Lturb and lower uair

' , autoignition was observed relatively further upstream with the SPP.

Autoignition delay timeEffect of air velocity. The effect of air velocity on autoignition delay time for the LPP and SPP is presented in Figure 5. Over the conditions presented, for the LPP, τ5 pc ranges from 15 ms to 35 ms and decreased with increasing U air at an average rate of 9.3 ms per 10 m/s U air. With the SPP, τ5 pc also decreased with increasing Uair.

As discussed in the previous section, L5 pc increased with increasing Uair, as observed by Markides et al. [30], and also Markides [19] for pre-vaporised n-heptane. However, rather than τ5 pc decreasing with increasing Uair, Markides [19] observed longer residence times at

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(a) (b)Figure 4. Effect of Uair on L5 pc at U fuel = 0.7 m/s (a) and 1.1 m/s (b) at six sets of constant T air conditions for two grid sizes, LPP M = 3 mm in empty markers and dash-dot lines, SPP M  

= 1.2 mm in filled markers and solid lines.

(a) (b)Figure 5. Effect of Uair on τ5 pc at U fuel = 0.7 m/s (a) and 1.1 m/s (b) at six sets of constant T air conditions for two grid sizes, LPP M = 3 mm in empty markers and dash-dot lines, SPP M =

1.2 mm in filled markers and solid lines

higher values of Uair. The different result with liquid fuels could be due to the evaporation of the fuel droplets. Schroll et al. [31] used three-dimensional direct numerical simulations to demonstrate the impact of evaporation of n-heptane droplets on the autoignition process under isotropic decaying turbulence. They found autoignition to be delayed with increasing initial droplet size. It has already been observed in this study that the droplet dmean decreased with increasing U air. These smaller droplets had a higher total surface area, resulting in faster generation of fuel vapours. Then, the decrease in τ5 pc with increasing Uair as observed in (b) appears to agree with the findings of Schroll et al. [31].

Effect of perforation size. Figures 5 and 6 compare τ5 pc between experiments with the LPP and SPP with the latter figure presented in the form of an Arrhenius plot. The results for both the LPP and the SPP show non-linearities in the Arrhenius plots, indicating that the autoignition reaction rate is not just kinetically-controlled but is affected by the flow conditions. Similar behaviour has been observed in autoignition of gaseous fuels [3,19].

For similar flow conditions, τ5 pc was shorter with the SPP than with the LPP, but this effect was more significant at the higher U fuel. This was in contrast to the work of Markides & Mastorakos [4], who observed an increased τ5 pc with a smaller grid size for a given Uair when studying gas fuel autoignition. The shorter τ5 pc with the SPP in this case could again be understood in terms of higher fuel evaporation rates due to smaller droplet diameters. Additionally, the LPP and SPP had different flow turbulence characteristics; the SPP flow fields were characterised by a larger Lturb and lower uair

' . This suggests that, with the SPP, both flow turbulence and vaporisation effects contributed towards the shorter τ5 pc.

CONCLUSIONS

The effect of turbulence on the autoignition of liquid n-heptane jets injected axisymmetrically from a circular nozzle into a confined turbulent coflow of high-temperature air at atmospheric pressure has been investigated experimentally. Measurements have been performed with air temperatures in the range 1107 – 1138 K, bulk air velocities in the range 14 – 32 m/s and fuel injection velocities of 0.7 and 1.1 m/s. The primary atomisation of the liquid n-heptane jet was characterised through high-speed imaging and the earliest or most upstream autoignition

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(a) (b)Figure 6. Effect of T air on τ5 pc at U fuel = 0.7 m/s (a) and 1.1 m/s (b) at six sets of constant Uair conditions for two grid sizes, LPP M = 3 mm in empty markers and dash-dot lines, SPP M =

1.2 mm in filled markers and solid lines.

location found (the 5th percentile location L5 pc), and a corresponding delay time, τ5 pc, identified. Turbulence was manipulated by varying the bulk air velocity, Uair, and performing experiments with two different turbulence-generating grids with hole sizes of 1.2 and 3 mm, known as the small perforated plate (SPP) and large perforated plate (LPP), respectively.

The mean droplet diameter decreased as Uair increased with both plates. Also, L5 pc increased with increasing Uair , indicating a downstream shift of the autoignition regions. This downstream shift is in agreement with the findings reported for the autoignition of gaseous fuels. However, in contrast to gaseous fuel autoignition, τ5 pc decreased with increasing U air, which may be, at least in part, due to faster fuel vapour generation with the smaller droplets at higher U air, which is in accordance with the literature. Arrhenius-type plots displayed non-linearities for both plates, indicating the finite effect of phase inhomogeneities and turbulence.

The mean droplet diameter was smaller, and the L5 pc shorter, when using the SPP compared to the LPP, although the latter effect was more pronounced at higher fuel velocities and air temperatures. The τ5 pc was also shorter with the SPP than the LPP, a result in contrast to what has been observed for autoignition of gaseous fuels. It is expected that both flow turbulence and vaporisation effects will contribute towards this result.

The autoignition of liquid fuels is a multiphase, multiprocess phenomenon and the effects of atomisation, evaporation, turbulent mixing and reaction can be coupled and must all be considered in each study. The results of this work can be used to understand the autoignition behaviour in the particular reactor of interest, which can be extended to similar situations, and can be used to develop and validate computational models of spray formation and non-premixed two-phase turbulent combustion. The understanding of liquid fuel autoignition can help develop cleaner and more efficient processes for combustion of fossil fuels.

NOMENCLATURE

d inj Inside diameter of fuel injector nozzle [m]Dinj Outside diameter of fuel injector nozzle [m]dqzr Inside diameter of quartz reactor [m]Dqzr Outside diameter of quartz reactor [m]L5 pc Most upstream autoignition location [m]Lturb Longitudinal integral turbulent lengthscale [m]M Turbulence generating perforated plate hole diameter [mm]

Qfuel Volumetric flow rate of liquid fuel [m3/s]T air Bulk air inlet temperature [K]T fuel Fuel temperature at injection [K]u' Velocity fluctuation [m/s]

⟨ uz(r )⟩ Local mean axial velocity [m/s]U air Bulk air inlet velocity [m/s]U fuel Liquid fuel injection velocity [m/s]

z Axial direction coordinate [m]

ρair Air density at T air and 1 atm pressure [kg/m3]

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ρ fuelamb Fuel density at ambient temperature (296 K) [kg/m3]

ρ fuel Fuel density at T fuel [kg/m3]τ5 pc Autoignition delay time [s]

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