In-nozzle flash boiling flow of multi-component fuel and its effect on near-nozzle spray

Fuel 252 (2019) 55–67

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Full Length Article

In-nozzle flash boiling flow of multi-component fuel and its effect on nearnozzle spray Shangze Yanga, Xuesong Lia, David L.S. Hunga,b, Masataka Araic, Min Xua,

T

⁎

a

School of Mechanical Engineering, Shanghai Jiao Tong University, Shanghai, China University of Michigan-Shanghai Jiao Tong University Joint Institute, Shanghai Jiao Tong University, Shanghai, China c Department of Mechanical Engineering, Tokyo Denki University, Japan b

A R T I C LE I N FO

A B S T R A C T

Keywords: Multi-component fuel Flash boiling spray Two-dimensional transparent nozzle Bubble flow Gasoline surrogate

In comparison with single-component fuel spray, the atomization and evaporation of multi-component fuel spray is more complex due to the physical property difference among the fuel components. This complexity is more significant when the fuel spray is formed under flash boiling condition. Therefore, the investigation of atomization and evaporation of multi-component fuel has recently become a practical research focus. However, due to the lack of related experiments and theoretical explanations, the spray formation under flash boiling condition has yet to be fully revealed. To address such challenges, this study is focused on high-speed optical investigation of multi-component flash boiling spray phenomena. Test fluids were gasoline surrogate fuels composed of three components (n-pentane, iso-octane, and n-decane). A scaled-up transparent two-dimensional nozzle was fabricated for the optical investigation. The droplet diameter was also measured using Phase Doppler Interferometry (PDI). Experimental results show that mixing of high volatility fuel with low volatility fuel can significantly enhance the generation of flash-boiling bubbles in the internal flow inside the nozzle (in-nozzle flow). Meanwhile, the increase of flash boiling bubbles can promote fuel breakup in the near-nozzle region, resulting in fuel droplets of smaller SMD with more uniform distributions. Besides, proper mixing ratio of high volatility fuel and low volatility fuel can moderate the layered structure (bubble and liquid flows) inside the nozzle and therefore can stabilize the breakup process from liquid ligaments into fuel spray with low variation of spray geometry.

1. Introduction Both improving fuel efficiency and reducing combustion emission have been considered as key goals for driving the future development of SI combustion engine. One promising way to achieve such goals is to actively control the atomization process. Flash boiling spray, in which the liquid fuel is partially or completely vaporized within an injector nozzle because the saturated vaporization pressure of the fuel is higher than the local pressure inside of the nozzle, has been considered as a potential solution for active spray control of practical multi-component fuel engines. In comparison with single-component fuel, the physical and chemical properties of multi-component fuel mixture are co-determined by the fuel components within, as well as their proportion, in the mixture. Under the flash boiling temperature/pressure conditions, different volatile levels of fuel components will induce different fuel evaporation characteristics. It will complicate further the atomization and evaporation processes of the mixture. Previous researches [1–7]

⁎

have revealed that mixing high volatility fuels into low volatility fuels enhanced atomization efficiency compared with the atomization of low volatility fuels only, and the emission related performance (such as NOx and total hydrocarbon (THC) emissions) could be improved to some extent. To illustrate the effect of component proportions on the performance of the multi-component fuel spray, numerous investigations have been conducted in recent years. For instance, Senda et al. [8] used a rapid compression expansion machine (RCEM) to investigate the characteristics of multi-component spray under flash-boiling conditions. Non-intrusive optical diagnostics methods such as laser-induced fluorescence (LIF) and Mie scattering were adopted to analyze the spatial distribution of liquid and gas phases of flash boiling spray. Their results showed that in the spray formation region located just outside of the nozzle, fuel components with higher volatilities mainly located at the upstream part of the spray jet near the nozzle exit, and fuel components with lower volatilities located at the downstream of the jet.

Corresponding author. E-mail address: [email protected] (M. Xu).

https://doi.org/10.1016/j.fuel.2019.04.104 Received 29 January 2019; Received in revised form 4 April 2019; Accepted 17 April 2019 0016-2361/ © 2019 Elsevier Ltd. All rights reserved.

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Accumulator Flow meter Heat exchanger

Magnetic valve

Heat source

2D Transparent nozzle

Xenon lamp

Precision slit Optical glass

High speed camera

Heater Synchronizer

Chamber

Signal Heat Fuel Gas Nitrogen bottle

Vacuum pump

Computer

Fig. 1. Experimental setup for shadowgraph measurement.

of fuel jet. However, this model was limited since it was assumed all flash boiling bubbles were of the same size and distributed uniformly within the nozzle, which was obviously quite inaccurate. Vidal et al. [25] simulated the formation of cavitation within the nozzle for multicomponent fuels. This simulation work was not fully completed since the cavitation bubble dynamics and the corresponding bubble eruption due to the liquid breakup were not properly modeled. Also, experimental validations were not included. To address the aforementioned issues, in this investigation, a twodimensional transparent nozzle was used for the optical investigation of multi-component fuel sprays under flash boiling conditions. Test fuels were gasoline surrogate fuels which were composed of three components (n-pentane, iso-octane and n-decane) in various proportions and the proportion was decided by the distillation experiment for each fuel. The in-nozzle flow characteristics and its response on the external characteristics of flash boiling sprays near nozzle field were analyzed under different test conditions. Phase Doppler Interferometry (PDI) was used to measure near-nozzle field droplet diameter characteristics. With these measurements, this investigation aims to compare the difference between single-component flash boiling sprays and multi-component flash boiling sprays, and to provide experimental data for numerical models of multi-component fuel sprays. The remaining description of this paper is organized as follows: Section 2 of this paper introduces the experimental setup and the data processing schemes. Distillation curves and other properties of multi component surrogate gasolines are discussed. Section 3 illustrates the experimental data concerning in-nozzle bubble flow and flash boiling spray. Finally, in Section 4, new findings are summarized as the conclusion of this research work.

Thus, the breakup process which mainly took place at the downstream of the liquid jet was dominated by fuel components with lower volatilities. Gemci et al. [9] mixed n-butane into hexadecane with different mixing proportions. Both the spray performance and droplet diameter distribution near the nozzle exit region were measured under different temperature conditions. It was shown that as the increase of the fuel temperature and the mass fraction of n-butane, the diameter of spray droplets decreased and the atomization efficiency of the mixture was improved. Researchers [10,11] adopted the fuel dependent laser dopant for each component of the fuel mixture and used LIF to evaluate the evaporation properties of each fuel component in the mixture. It was shown that sequential evaporation took place among different fuel components, which had a positive impact on spray ignition and combustion stability. Zigan et al. [12] found that although the addition of heavy fuel components deteriorated the breakup performance compared with the liquid jet with light fuel only, the variation of the fuel spray variation was reduced by such addition. Machado et al. [6] utilized an analogical method of a triangular coordinate system to analyze the impact of every fuel composition of a three-component fuel mixture on the morphology of spray characteristics such as the spray angle and jet length, and related combustion efficiency. Despite of the research works previously introduced, the impact mechanisms of multi-component fuel compositions on external spray characteristics are still not clear. Meanwhile, it has been demonstrated that the multi-phase dynamics of the in-nozzle flow has significant impact on the external characteristics of sprays under flash boiling or cavitating flow conditions [13–19]. The illustration of component proportions influence on the multi-component fuel in-nozzle flow is crucial before understanding the external performance of sprays, especially for flash boiling sprays [20–22]. However, such kind of investigations, although significant, are lacking in existing literatures. On the other hand, in absence of such kind of experimental data discussed above, existing numerical models for flash boiling sprays of multi-component fuel are usually based on mostly theoretical assumptions rather than experimental measurements. For example, the numerical models in [23,24] did not incorporate neither the modeling of in-nozzle flows nor detailed spray breakup process. The model in literature [3] considered the impact of in-nozzle bubbles on the breakup

2. Experimental setup and image post processing procedures 2.1. Experimental scheme In this investigation, we implemented both shadowgraph and PDI measurements for investigating the in-nozzle flow performance and external spray characteristics. Fig. 1 depicts the experimental setup of the shadowgraph measurements. A constant volume chamber filled with nitrogen was used for the experiments conducted in this work. The 56

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thickness direction, which would improve the measurement quality when conducting the shadowgraph experiments. And the occurrence of cavitation was restricted by the thin 2D nozzle. As shown in Figs. 7 and 8, limited cavitation could be found in the inlet corners under the subcooled test conditions. It is worth noting that with the increased of fuel temperature, cavitation was captured in the inlet corners of the nozzle and the presence of cavitation can alter the flashing behavior in the downstream. However, these two phenomena are almost impossible to separate under flash boiling test conditions. Nevertheless, it should be noted that the vaporization in either phenomenon was caused by a greater saturated pressure than the static local pressure, and it is not the goal of this work to distinguish the two phenomena. In this work, we tend to treat these two phenomena as a whole part to analysis the influence of proportion of multi-component fuel on the fluid performance. A scaled-up nozzle was used with the nozzle width of 2 mm and nozzle length of 10 mm. The use of such a scaled-up nozzle can help the analysis of the in-nozzle flow by offering more flow details. The design of this two-dimensional transparent nozzle is depicted in Fig. 2. It is worth noting that results from the two-dimensional nozzle could not be applied to practical engine applications directly, while it provides some crucial fundamental aspects in interpreting the physics behind flash boiling process. Such observation is challenging to be captured if using practical injector or three-dimensional transparent nozzle. Otherwise, we have used boundary conditions that are correlated to actual engine conditions. Fig. 3 shows the experimental setup of PDI droplet diameter measurement. PDI is used to quantify the impact of component proportion of multi-component fuel on the droplet diameter. In this research, an Artium PDI-x00MD system was used to measure the spray droplet size distribution over a specific sampling time. Most of the setup for generating flash boiling spray was similar to the shadowgraph experiment. The transmitter and receiver of the PDI device were placed on the same horizontal plane and their positions were adjusted to measure droplet diameter at different locations. The intersection angle between the transmitter and receiver was set to 150° and the width of the aperture was set to 50 μm for an optimal measurement quality. Droplets with a diameter greater than 100 μm were filtered out to remove interference or noise. A total of 9000 drop counts was measured at each point to

Fig. 2. Schematic of the transparent two-dimensional nozzle.

internal pressure of this chamber could be fully adjusted by a vacuum pump which was connected to it. The control of fuel temperature was regulated by a heat exchanger in the upstream of the two-dimensional transparent nozzle. To minimize the heat loss as the fuel flowed in and out across the nozzle, an electrical heater was implemented in the chamber to control the temperature of the ambient gas (nitrogen) so that the fuel temperature and the ambient temperature were the same. Another pressure adjustable nitrogen source was used to pressurize a fuel accumulator and maintain an injection pressure of 5.0 MPa. A high precision servo valve (magnetic valve) was installed to control the injection duration for each experiment. In this investigation, the injection duration was fixed at 10 ms to deliver a stable spray structure during the injection. A fuel filter was also installed with the mesh size of 5 µm to filter out any particle debris which might affect the nucleation in the nozzle. A high-speed camera (Phantom V1210) was operated at 30 kHz to record the bubble behavior and atomization phenomena. A highpower Xenon lamp (Newport, 300 W continuous) was used as the light source for shadowgraph measurements. All experimental apparatus was synchronized by a time delay generation device. Fig. 2 depicts the design of the two-dimensional transparent nozzle as circled by red dash line in Fig. 1. The two-dimensional transparent nozzle was formed by a thin precision gasket clamped by two optical glass blocks. The thickness of the precision gasket used in this research was 40 μm to reduce the overlaps of in-nozzle bubbles along the

Flow meter Heat exchanger

Heat source

Accumulator 2D Transparent nozzle

Transmitter

Chamber Heater Synchronizer Signal Heat Fuel Gas Laser Nitrogen bottle

PDI Signal processor Fig. 3. Experimental scheme with the PDI system incorporated. 57

Computer

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Boiling point: Decane (174.1 )

BP: iso-Octane (99 )

Boiling point: n-Pentane (36.1 )

Fig. 4. Distillation curves for the three surrogates used in this research.

Fig. 5. Physical properties of fuel surrogates as a function of fuel temperature. Panel (a): Fuel bubble point pressure/vapor pressure, Panel (b): Fuel density, Panel (c): Fuel viscosity, and Panel (d): Fuel surface tension. 58

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T=100

T=112

T=122

T=132 T=143

Bubble extraction

Region1

10mm

T=92

X

1mm

Nozzle inlet

Nozzle outlet

12mm

2mm Region2

Y

Jet length

2mm

2mm

-3 -2 -1 0 1 2 3

15mm

(a) Ca=0.97 Ca=0.96 Ca=0.94 Ca=0.92 Ca=0.90 Ca=0.87 Re=9500 Re=9800 Re=10600 Re=11200 Re=11900 Re=12600 We=6700 We=7000 We=7700 We=8400 We=9200 We=10000

T=92

PDI test position

T=100

T=112

T=122

T=132 T=143

18mm

Raw data

Jet length extraction

Fig. 6. Data post processing and definitions used in this work.

T=66

T=76

T=84

T=92

T=100

(b)

1mm

T=59

Ca=0.98 Ca=0.98 Ca=0.96 Ca=0.95 Ca=0.93 Ca=0.92 Re=8300 Re=8900 Re=9500 Re=10100 Re=10700 Re=11400 We=6300 We=6600 We=7200 We=7700 We=8300 We=8900

T=92

T=100

T=112

T=122

T=132 T=143

(a) Ca=0.98 Ca=0.97 Ca=0.95 Ca=0.94 Ca=0.92 Ca=0.90 Re=11100 Re=11600 Re=12300 Re=12900 Re=13500 Re=14100 We=6200 We=6500 We=7000 We=7600 We=8100 We=8800

T=92

T=100

T=112

T=122

T=132 T=143

(c) Ca=0.99 Ca=0.99 Ca=0.98 Ca=0.96 Ca=0.95 Ca=0.94 Re=7700 Re=8100 Re=8800 Re=9400 Re=9900 Re=10600 We=6100 We=6300 We=6800 We=7400 We=7700 We=8400 Fig. 8. In-nozzle flow and issuing jet morphologies for multi-component fuels from 92 °C to 143 °C. Panel (a): Surrogate 1, Panel (b): Surrogate 2, and Panel (c): Surrogate 3.

(b) Table 1 Fuel composition of the three surrogates used.

Ca=1.02 Ca=1.02 Ca=1.02 Ca=1.01 Ca=1.01 Ca=1.01 Re=5400 Re=5800 Re=6300 Re=6800 Re=7300 Re=7800 We=4900 We=5000 We=5300 We=5500 We=5800 We=6100 Fig. 7. In-nozzle flow and issuing jet morphologies for single component fuels under various temperature conditions. Panel (a): n-pentane in the temperature range from 59 °C to 100 °C. Panel (b): n-decane in the temperature range from 92 °C to 143 °C.

make sure a statistically meaningful data collection.

Items

Volume fraction of MCFs (n-Pentane:iso-Octane:n-Decane)

Surrogate 1 Surrogate 2 Surrogate 3

0.37:0.48:0.15 0.21:0.58:0.21 0.14:0.54:0.32

octane, and n-decane. Three different types of fuel surrogates, namely surrogate 1, surrogate 2, and surrogate 3, were prepared to mimic the gasoline fuel and their fuel composition can be found in Table 1. Fig. 4 depicts the distillation curves of each fuel surrogate used in this work. As shown in Table 1 and Fig. 4, for the surrogate 2, the proportion was

2.2. Test fuels and fuel surrogates The component fuels used for fuel surrogates were n-pentane, iso59

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pressure of the ambient. Ps is fuel saturation vapor pressure. The ratio of the ambient pressure Pa and fuel saturation vapor pressure Ps (bubble point pressure for multi-component fuels) is used to define the superheat degree (SD, Pa/Ps or Pa/Pb). We have summarized all the experimental fuel temperatures and their corresponding superheat degree in Table 3. It is worth noting that changing ambient pressures also could be used to control the superheat degree, as well as many other factors. In this work, we will limit our discussion on the impact of superheat degrees via varying fuel temperatures only. For PDI measurements, three temperatures (122 °C, 132 °C and 143 °C) were chosen because reasonable breakup phenomenon was observed for these conditions and no heavy dense region of spray existed under these conditions, which might deteriorate the quality of PDI measurements.

adjusted to approximate its distillation behavior with that of a practical gasoline fuel (Ron 98). As for surrogate 1, the proportion of high volatility fuel (n-pentane) was increased such that its distillation curve was consistently below the surrogate 2. Similarly, as the proportion of low volatility fuels (n-decane) increased, the distillation curve of surrogate 3 was raised above that of the surrogate 2. The boiling temperatures of the three single component fuels were also plotted in Fig. 4. As seen, the distillation curves of the three surrogates were all within the regime formed by the boiling temperature of n-pentane (lowest) and n-decane (highest). The flash boiling performance of the sprays is impacted by various physical properties of the fuel surrogate, such as the saturation temperature/pressure, density, dynamic viscosity, and surface tension. It is worth noting that these properties are for the liquid phase of the surrogates. The physical properties of the fuels were obtained from NIST database. Based on the data for each single fuel component, we calculated these physical properties from their fuel composition based on the REFPROP model [26]. These properties as a function of fuel temperature were summarized in Panel (a) to Panel (d) of Fig. 5. It is worth mentioning that we used the bubble point pressure as in Panel (a) of Fig. 5 to demonstrate the vaporization behavior of the fuel surrogate, which is closely connected to the flash boiling phenomenon. It is such defined that at the bubble point pressure, the first bubble of vapor is formed as the surrogate temperature increases. As seen in Fig. 5, the differences among the physical properties of the fuel surrogates are not significant. For all other properties analyzed, their function curves vs. temperature are similar and close to the property of iso-octane. When a higher portion of high volatility fuel is mixed into the low volatility fuel, the bubble point pressure increases, while its density, dynamic viscosity, and surface tension decrease. From these estimations, it can be seen that the surrogates have similar overall physical properties. Therefore, any performance variation in the flash boiling process, if significant, should not be attributed to fluid dynamics reasons, but different evaporation/boiling characteristics of each single component fuel in the surrogate.

2.4. Data post processing procedures Fig. 6 depicts the post-processing scheme used in this work and some of the definitions which are useful for the analysis. On the lefthand-side, a representative raw measurement is shown, which demonstrates the transient states of both in-nozzle flow and near-nozzle spray. The bubble formation for in-nozzle flash boiling flow is caused by the heterogeneous nucleation [16], and the bubbles are formed on both sides of the nozzle wall. Shadowgraph technique can capture the variation of fluid distribution, and it is clearly evident to infer that the dark region in the nozzle represents the bubble/gas phase, and the dark region outside of the nozzle represents the spray/liquid phase. For the analysis carried out in the following sections, we define the direction that along the internal flow and spray flow as Y-direction and the transverse direction as X-direction. A binarization procedure was implemented to calculate the bubble fraction (area of bubbles divided by the total area of the nozzle) in region 1 and to define the jet length (as shown in the right-hand-side of Fig. 6). For the in-nozzle section, we set the cutoff value to be 60% of the maximum pixel value in this section. For spray liquid jet length extraction, the cutoff value was set to be 20% of the maximum pixel value and any pixel value below the thresholds was considered liquid jet, and the threshold values were chosen based on [28]. The jet length represents the primary breakup length of the flash boiling spray. The interface between the liquid phase and the gas phase is plotted as seen in Fig. 6. A 2 mm by 2 mm square region (region 2) just beneath the nozzle exit is chosen to illustrate the fluctuation of the spray external characteristics. The 2 mm distance in y direction was chosen arbitrarily, which can capture the most fluctuating feature in the nearfield. A correlation coefficient Cij is used to quantify the spray fluctuation in region 2, which can be expressed by:

2.3. Test conditions In this work, a constant ambient pressure of 1 bar was used for all tests. The fuel temperatures were selected when 30%, 50%, and 70% volume percentage of the fuel surrogates were vaporized. As shown in Fig. 4, different temperatures were used in the range from 59 °C to 143 °C. The test conditions were summarized in Table 2. Noted that we also provided estimated values for Reynolds Number, Weber Number and Cavitation Number for the experiments, as shown in Figs. 7 and 8. Reynolds Number and Weber number was estimated by the mean velocity which was calculated from the flow rate. Cavitation Numbers was defined as a ratio of pressure differences [27], as seen in Eq. (1).

Ca =

−

(∑

m

where Pinj is the injection pressure upstream and Pback is the back Table 2 Test Conditions. Specifications

Fuel type

n-pentane, n-decane, surrogate 1, surrogate 2, surrogate 3 5 101 59–143 10 5400–14,100 4900–10,000 0.87–1.02

Injection pressure (MPa) Ambient pressure (kPa) Fuel temperature (°C) Injection duration (ms) Reynolds Number (Re) Weber Number (We) Cavitation Number (Ca)

−

)

j

−

∑n (Fimn − F )2 ⎛∑m ∑n (F jmn − F )2⎞ j i ⎠ ⎝ ⎜

⎟

(2)

where Fi and Fj are two consecutive scalar matrices representing the intensity of the images (for region 2), m is the number of rows of the image, n is the number of columns of the image, and mn represents a − pixel in the image defined by m and n. F is the averaged intensity for this specific image. From the definition of Eq. (2), the closer Cij to 1.0, the stronger is the correlation between the two consecutive images. Otherwise, the correlation between the two images is weaker as Cij decreases, indicating a stronger fluctuation between the two images. For PDI measurements, we chose three distances of 12 mm, 15 mm, and 18 mm away from the nozzle exit in the Y direction. The spatial resolution in the X-direction was set to 1 mm, and 7 data points were collected for each distance (a total of 3 distances). A total of 21 data points was used for each injection condition. Then, the Sauter mean diameter (SMD) was calculated by Eq. (3) for further analysis:

(1)

Items

i

Cij =

pinj − ps pinj − pback

−

∑m ∑n (Fimn − F )(F jmn − F )

SMD =

∑ Dn3/∑ Dn2 n

60

n

(3)

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Table 3 Superheat degree. Tfuel Fuel

59 °C

66 °C

76 °C

84 °C

92 °C

100 °C

112 °C

122 °C

132 °C

143 °C

n-Pentane n-Decane Surrogate 1 Surrogate 2 Surrogate 3

0.48 69 0.88 1.2 1.6

0.39 48 0.72 1.0 1.3

0.30 29 0.55 0.78 1

0.25 21 0.45 0.64 0.81

0.20 15 0.37 0.52 0.67

0.17 11 0.31 0.43 0.55

0.13 6.4 0.24 0.33 0.42

0.11 4.6 0.20 0.27 0.34

0.09 3.3 0.16 0.22 0.28

0.07 2.3 0.13 0.18 0.23

surrogate 1, surrogate 2, and surrogate 3, respectively. To compare the impact of fuel composition on the morphology of in-nozzle bubble formation and emerging jet, the shadowgraphs at the same temperatures from 92 to 143 °C are demonstrated. From Fig. 8, it can be seen that as the fuel temperature increased, the volume of flash boiling bubble generated for all the three fuel surrogates, and the increase of flash boiling bubble fraction had a direct impact on the breakup and atomization of near-nozzle liquid jet. Meanwhile, as the percentage of the low volatility fuel increased, the component proportion changes would lead to the suppression of flash-boiling bubble formation (since the percentage of fuel that is hard to vaporize increases), which in turn deteriorated the atomization of the liquid fuel. Similar to the discussions in Fig. 7, the initial fuel temperature with visible flash boiling bubbles was 112 °C for the surrogate 3, as compared with 76 °C for npentane single component fuel. Even though 14% volume of n-pentane in the surrogate 3 would theoretically evaporate and form flash boiling bubbles, such evaporation process was not notable for fuels with a temperature lower than 112 °C. Meanwhile, with the addition of high volatility fuel components into n-decane (with no flash boiling bubbles seen until the fuel temperature reached a temperature of 143 °C), flash boiling bubbles appeared before the fuel temperature reached the boiling point of n-decane. As the temperature of the fuel mixture reached the boiling point of high volatile components, the bubble nucleation of such fuel would progress with the size of the nucleus too small to be captured by typical morphology measurements. The change of fuel physical properties, including surface tension, viscosity, and density, might also affect the growth process of the flash boiling bubbles. Furthermore, the growth of nucleus bubbles would depend on sufficient heat transfer and high volatile fuel mass transfer to them. If insufficient energy (fuel temperature) or high volatility fuels (fuel concentration) could be provided, no visible flash-boiling bubble would be seen in the nozzle, while the impact of such potential nucleation could still be observed on the fuel (such as surrogate 2 under the temperature of 100 °C, as shown in Panel (b) of Fig. 8.). Finally, as seen in Panel (a) of Fig. 8, flash boiling bubbles started to form under the fuel temperature of 92 °C, in which the vaporization of iso-octane and n-decane was likely to be impossible. The local pressure in the nozzle would be higher than 1 bar, indicating the local boiling temperature for iso-octane and n-decane would be even higher than 100 °C (compared with the distillation curves in Fig. 4). Although it would be quite challenging to obtain direct experimental evidence to show the composition of the gas phase within the flow channel, we suspect that fuels with higher volatilities (such as n-pentane) would evaporate first. The flash boiling bubbles would affect both the external fuel jet breakup and the combustion process since fuels with lower volatiles would tend to locate in the center region of the spray. With the shadowgraph images, we used the post-processing method introduced in Fig. 6 to interpret the flash boiling characteristics quantitatively by means of the bubble fraction within the nozzle and the jet length in the spray for multi-component fuels. Results as a function of fuel temperatures are incorporated in Fig. 9 with Panel (a) and Panel (b) summarizing the bubble fraction and jet length, respectively. Here, the bubble fraction represents the average void (bubble) fraction in region 1. As shown in Panel (a) of Fig. 9, as the concentration of high volatility fuels increased, the bubble fractions in the nozzle also

It is worth noting that the data of droplet size using to calculate the SMD has been normalized based on the spray concentration and the capture frequency. We sampled 1800 droplets for each point and 4 repeats were carried out for the uncertainty analysis purpose. As for cause and effect relationship between flash boiling and spray formation, bubble fraction and layered structure explained in Section 3.2 are the representative characteristics of region 1 where physical causes for atomization promotion are prepared. The jet length, SMD of measured spray, and correlation coefficient in region 2 are the resulting effects of these causes. In the following sections, the causal relationship between these phenomena are discussed by showing the experimental data and illustrating the physical mechanisms of atomization. 3. Result and analysis 3.1. Impact of fuel composition on the morphology of In-nozzle flow and external spray Flash boiling sprays of single component fuels were observed before we studied the performance of flash boiling sprays using the fuel surrogates. Fig. 7 depicts the in-nozzle flows and emerging liquid jets for npentane and n-decane, respectively. Panel (a) of Fig. 7 shows the shadowgraphs of n-pentane sprays under the fuel temperature from 59 °C to 100 °C. As seen from the photographs, when the fuel temperature was at 59 °C, no flash boiling bubble was seen inside the nozzle. The innozzle flow was in liquid phase and a wide external liquid jet was evident, with a negligible amount of liquid jet breakup outside of the nozzle. The liquid jet converged slightly as Y increased, but not significantly. As the temperature was elevated to 66 °C, a notable change took place outside of the nozzle as no obvious change was seen in the nozzle. The liquid jet converged rapidly as the jet flowed away from the nozzle exit. We suspect that such phenomena were a result of the existence of tiny bubbles within the nozzle that could not be properly captured by the spatial resolution of the camera used in the experiments. However, visible fuel droplets could be identified near the liquid jet, indicating an improved jet breakup tendency. As the fuel temperature further increased from 66 °C to 100 °C, visible flash boiling bubbles could be captured within the nozzle, causing a more pronounced flash boiling induced breakup and flow rate decrement, as described previously in [18]. An increase of in-nozzle bubble fraction also shortened the jet length with the impact of flash-boiling. Panel (b) of Fig. 7 shows the spray characteristics with n-decane. Fuel temperature was selected between from 92 °C and 143 °C. Since the fuel temperature was still lower than the boiling temperature of n-decane (174.1 °C), no flash boiling phenomenon was observed in this fuel temperature range. For the n-decane injection, as the fuel temperature increased, the jet width increased as well because of the decrease of fuel surface tension, viscosity, and density. Thus, we can see that the converging feature of the n-pentane jet at 66 °C was mainly due to other issues such as phase change rather than the variation of fluid dynamic properties. It is also worth noting that as seen in Panel (a), when flash boiling bubbles took place in the nozzle, a distinctive layered (or wavy) structure of the flash boiling bubbles could be clearly seen in the nozzle. Such structure might affect the stability of the emerging liquid jet and we will discuss such features in next sections. In comparison, Fig. 8 demonstrates similar observations for 61

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Fig. 9. In-nozzle bubble fraction and jet length as a function of temperature for different multi-component fuels. Panel (a): Bubble fraction as a function of fuel temperature and Panel (b): Jet length as a function of fuel temperature.

With the bubble point pressure increases from 0.2 MPa to 0.5 MPa, the bubble fraction increases by 35% and the jet length decreases by 80%. Therefore, the bubble point pressure of the fuel mixture can be used as a flash boiling indicator with respect to the bubble fraction and jet length of emerging liquid jet.

increased notably. Meanwhile, the bubble fraction trends for surrogate 2 and surrogate 3 were similar when compared to that of surrogate 1. It is because surrogate 1 had a substantially higher concentration of npentane that yielded evaporation at lower temperatures. As the fuel temperature further increased to 143 °C, the bubble fraction for surrogate 2 increased sharply because of likely the nucleation of iso-octane under this condition. For the surrogate 2, a 10% higher concentration of n-pentane and iso-octane had induced an 8% higher bubble fraction at the temperature of 143 °C. Correspondingly, the jet length was found to be strongly related to the bubble fraction within the nozzle. As the fuel temperature and the bubble fraction increased, the jet length decreased. As shown in Panel (b) in Fig. 7, The jet length for surrogate 1 was the shortest and the liquid jet almost diminished under the fuel temperature of 143 °C. Under this fuel temperature, both surrogate 2 and surrogate 3 demonstrated a reasonable breakup efficiency. As to surrogate 1, a 10% higher bubble fraction induced 15% shorter jet length under the fuel temperature of 143 °C. As compared with Panel (b) of Fig. 7 with pure ndecane as the fuel, no visible atomization was observed. It is clear that the atomization of the three surrogates was dominated by the thermodynamic process of the fuel (flash boiling) rather than liquid breakup induced by fluid dynamics. With the analysis in Fig. 9, it is worth noting that the initiation and growth of flash boiling bubbles are mainly determined by the pressure gradient between the nucleus and the surrounding liquid phase. For single component fuels, this pressure difference can be approximated by the saturation vapor pressure and the local pressure near the bubble. However, for multi-component fuels, there is no well-accepted definition of the saturation vapor pressure for a specific fuel composition [8]. Instead of the saturation pressure, the bubble point pressure is used which comes from both theoretical estimations and empirical experimental data. Fig. 10 depicts the bubble fraction and jet length for the multicomponent fuels again, but as a function of the bubble point pressure of the fuel (see Panel (a) in Fig. 4). The bubble point pressure varies as a function of fuel temperature and fuel composition, which can be calculated by the NIST database. As seen in Panel (a) of Fig. 10, the difference of bubble fraction under the same bubble points was much smaller compared with those under the same fuel temperature. Very similar observations can be made for the jet length under the same bubble point pressure as well. The jet lengths for all three surrogates dropped steadily between the bubble point pressures of 0.2 MPa and 0.5 MPa, then the jet length decreased slowly until the liquid core completely disappeared by the impact of flash boiling bubble eruption.

3.2. Impact of fuel composition on the stability of spray formation As we discussed in the previous section, the bubble fraction in the nozzle increased with an increasing fuel temperature, and the value of bubble fraction tended to be similar under the same bubble point pressure for multi-component fuels. Meanwhile, as we discussed, the morphology, i.e., the profile of the bubble-liquid interface, turned out to be slightly different for single-component fuels and multi-component fuels. Fig. 11 shows the enlarged local characteristics of flash boiling bubble regions near the outlet of the nozzle. As shown in Panel (a) of Fig. 11, the bubble region for single component cases represented a “layered” structure at 92 °C fuel. It is hard to determine if the bubbles have jointed into a single bubble or it is just clouds of small bubbles, while we believe the impact of such consideration is less significant on the characteristics of the external spray. In comparison, the bubble structures and the contour of the gas-liquid interfaces in Panel (b)–(d) of Fig. 11 for multi-component fuels were smoother, and the layered pattern was obviously less significant. From our previous research, the characteristics of the in-nozzle bubbles would directly affect the dynamic of the sprays in the near-nozzle field [18]. Therefore, it is important to examine the impact of bubble structures on the sprays, even the in-nozzle bubble fractions in region 1 are similar. To quantitatively compare the fluctuation of sprays with different fuel compositions, we used the definitions set forth in Section 2.4 to calculate the correlation coefficient. As the flash boiling sprays were fully developed and stabilized, we chose 100 consecutive shadowgraph images to carry out the analysis in the near-nozzle field, and more specifically, in region 2 noted in Fig. 6. According to the definition, a higher correlation coefficient value indicates weaker fluctuations in the region. Fig. 12 depicts the correlation coefficient variation under different fuel temperatures for the fuels used in this study. Panel (a) to Panel (d) of Fig. 12 demonstrate the correlation curves for n-pentane under the fuel temperature of 76 °C, 84 °C, 92 °C, and 100 °C, respectively. For Panel (a), few flash boiling bubbles could be seen in the nozzle. In this case, atomization induced by flash boiling was not significant, yielding a higher overall correlation coefficient. As the temperature increased, the overall correlation coefficient dropped and the fluctuation of the 62

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(a)

(b)

Out of window

Fig. 10. In-nozzle bubble fraction and jet length as a function of bubble point pressure for different multi-component fuels. Panel (a): Bubble fraction as a function of bubble point pressure. Panel (b): Jet length as a function of bubble point pressure.

correlation curve was more pronounced. In Panel (b), the correlation coefficient oscillated around the value of 0.5, while in Panel (d) the averaged correlation coefficient was close to 0.2, indicating that a higher flash boiling level of the single component fuel could increase the bubble fraction, and causing strong oscillations of the external spray. However, it is potentially not desirable for practical fuel spray applications, for instance, excess oscillations might be the cause of damages associated with flash-boiling sprays. Panel (e) to Panel (h) of Fig. 12 depict the correlation curves for the multi-component fuel surrogates under the fuel temperature of 112 °C, 122 °C, 132 °C, and 143 °C, respectively. As seen, surrogate 3 had the most stable spray structure, followed by surrogate 2, with surrogate 1 at the end. With the increase of fuel temperature, the fluctuation of the multi-component fuel spray was stronger, but the stability of the nearnozzle field spray was still better than that with single component fuels. Although we have demonstrated that flash boiling sprays with more low volatility fuel in surrogate could achieve the higher stability, this performance might be caused by a higher boiling temperature and weaker flash boiling interaction for the low volatility fuel. Therefore, Fig. 13 illustrates similar comparisons of single component fuels and fuel surrogates, but the comparison was made under the similar saturation vapor pressure (for single component fuels) and bubble point pressure (for multi-component fuel surrogates). Under such conditions,

Fig. 11. Characteristics of flash boiling bubbles near the outlet of the nozzle. Panel (a): n-Pentane, fuel temperature of 92 °C. Panel (b): surrogate 1, fuel temperature of 132 °C. Panel (c): surrogate 2, fuel temperature of 132 °C. Panel (d): surrogate 3, fuel temperature of 132 °C.

(e)

(a)

Tf =112

Tf =76

(b)

(f)

Tf =84

Tf =122

(c)

(g)

Tf =92

Tf =132

(d)

(h)

Tf =100

Tf =143

Fig. 12. Spray oscillations represented by the correlation index of adjacent camera frames. The correlation coefficient was calculated by Eq. (2) introduced in Section 2.4. Panel (a) to Panel (d): Correlation coefficient as a function of injection time for pure n-pentane. Panel (e) to Panel (h): Correlation coefficient as a function of injection time for multi-component fuel surrogates. 63

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Fig. 13. Spray oscillations represented by the correlation index of adjacent camera images. The comparison was made under a similar saturated vapor pressure/bubble point pressure. Panel (a): Vapor pressure/bubble point pressure of 0.34 MPa, Panel (b): Vapor pressure/bubble point pressure of 0.41 MPa, and Panel (c): Vapor pressure/bubble point pressure of 0.49 MPa.

boiling bubbles, its concentration around the bubble drops significantly. While the supplement of light components is enabled by the multi-component diffusion across the components, which would also require a considerable time. As a result, the growth of the flash boiling bubbles would be suppressed, but meanwhile, the heat and mass transfer may happen simultaneously under this scenario, and it weakens the cyclical responses of the flash boiling bubbles and produces stable sprays. Therefore, the flash boiling behavior can be controlled by fuel composition modifications, and fuel with proper components or additives may realize superior flash boiling performance than typical fossil fuels. Furthermore, the growth of single component fuel bubbles could be illuminated by Rayleigh-Plesset equation [32] as Eq. (4):

in-nozzle bubble fractions for these cases were close to each other, so that the comparison of spray stability would be more focused on the structures of the bubbles instead of the volume of the in-nozzle bubbles (averaged bubble fraction in region 1). Panel (a) to Panel (c) of Fig. 13 depicts the comparison under the saturation vapor pressure/bubble point pressure of 0.34, 0.41, and 0.49 MPa, respectively. As seen from Fig. 13, even with a similar bubble fraction within the nozzle, the spray performance of single component fuels and fuel surrogates were quite different. For all three cases, the stability of surrogate 3 was the highest, followed by surrogate 2, surrogate 1, and finally single-component fuels. Therefore, multi-component fuels can indeed assist in generating more stabilized flash boiling sprays with roughly the same amount of bubbles within the nozzle. As discussed before, the layered, or wavy bubble, structures in the nozzle for single component fuels under the flash boiling conditions might be the root cause for the significant spray fluctuations. It is believed that the reason for these bubble features of single component fuels are caused by the mass and heat transfer between the bubble phase and the liquid phase. The energy source for the bubble nucleation of single component fuel is the internal energy of the surrounding liquid, and the source for mass transfer is also the surrounding liquid (noting that the surrounding liquid can evaporate into flash boiling bubbles if the temperature is sufficient). Therefore, heat transfer is the dominating term in the phase change process for single component fuels. As the nucleus bubble absorbs heat from the surrounding fuel, local temperature drops and the bubble growth would cease as the local temperature is lower than local boiling temperature. Then as heat is transferred to the surrounding liquid around the bubble, the heat transfer process may continue and then flash boiling bubble would continue to grow. Therefore, the cyclical bubble growth behavior shapes the cyclical layered structure of the flash boiling bubble aggregates, which consequently causes stronger fluctuations in the near field of the spray. As for multi-component fuels, previous research on the topic of boiling has demonstrated that the boiling of fuel mixture could indicate slower heat and mass transfer [29], slower nucleation process [30], and slower bubble growth rate [31]. The reason of these behaviors is that although the energy/heat transfer process of multi-component fuel is analogous to that for single component fuels, the mass transfer process is different. For instance, as the lighter components evaporate into flash

Pv − P∞ = ρl R

μ dR d 2R 3 dR 2 σ + ρl ⎛ ⎞ + 4 l +2 2 dt 2 ⎝ dt ⎠ R dt R

(4)

where Pv is the saturation pressure of the fuel, P∞ is the local pressure outside the fuel vapor bubble, R is the radius of the fuel vapor bubble, ρl is the density of the liquid fuel, μl the dynamic viscosity of the liquid fuel, and σ the surface tension of the liquid fuel. As seen in Eq. (4), the growth of flash boiling bubble is mainly under the function of the balance between pressure difference and surface tension. The force balance analysis is based on an assumption that the liquid at the boundary has constant density and viscosity. However, as to the multi-component fuel bubble, the complexity of physical properties of the mixture would influence the force balance analytical process and physical properties of the fuel in the boundary change with the bubble growth. Besides, the mass transfer process between different components influences the mass conservation analytical process for the derivation of Rayleigh-Plesset equation. For the said reasons, mass transfer process and physical properties of the mixtures should be considered in the dynamic analysis of multi-components fuel bubble’s growth. These issues are beyond the scope of this current study and will be investigated in future research. 3.3. Spatial distribution of droplet size Fig. 14 shows the measured axial SMD results at different test planes (y = 12 mm, 15 mm, and 18 mm) to quantify the atomization efficiency of multi-component flash boiling sprays. It is also of some value to 64

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(a)

Tf =122

(b)

Tf =132

(c)

Tf =143

Fig. 14. SMD as a function of axial position for multi-component fuel surrogates under different temperatures. Panel (a): Fuel temperature of 122 °C, Panel (b): Fuel temperature of 132 °C, and Panel (c): Fuel temperature of 143 °C.

radial locations. The results are presented in Fig. 15. We have previously discussed the impact of fuel temperature and fuel composition along the axial direction as shown in Fig. 14. Here we analyze these impacts in the radial direction. As seen in Panel (a) to Panel (c) in Fig. 15, with the fuel temperature of 122 °C, all fuel surrogates exhibited similar SMD patterns. The SMDs in the central locations were larger. At the outside regions of the spray, the SMDs were smaller because of the fuel-air interaction at these locations. More specifically, as the surrogate 1 droplets moved downstream, the SMD variation in the radial direction decreased. With a higher fuel temperature of 132 °C, similar trends were seen for surrogate 2 and surrogate 3. However, for surrogate 1, the droplet size was smaller in the central region but larger at outer locations. We suspect that the droplets at the outer region experienced more collisions, which increased the SMD of fuel droplets. For the fuel temperature of 143 °C, as seen in Panel (g) to Panel (i), the SMD distributions in the radial direction generally agree with our previous discussions. The droplet diameter in the central region for surrogate 3 was greater, while for surrogate 1 and surrogate 2, the droplet size in the central region was smaller. Furthermore, the impact of fuel composition on the near-filed droplet size distribution was analyzed in this study. The spray droplet number distribution from axial position of y = 15 mm under different fuel temperature was calculated and the results are presented in Fig. 16. As seen from Panel (a) to Panel (c) of Fig. 16, the fuel composition of multi-component surrogates has a strong impact on droplet size distribution. The droplet counts for surrogate 1 was peaked at a smaller droplet diameter, in the range of 10–20 μm. As the proportion of low volatility fuel increased, the distribution of droplet is more even across different droplet size regions and more droplets with greater diameters were captured. The reason for such observation is that the increment of low volatility fuel would suppress the formation of flash boiling bubbles, which in turn deteriorated the breakup of liquid fuel jet. Meanwhile, from Fig. 16, it can be seen that as the fuel temperature increased, a higher number of smaller droplets was captured for the same reason.

present the SMD results of single-component fuels, while such information is not strongly associated with this work and it could be found somewhere else in our previous publications [33]. The results are analyzed under different fuel temperatures, which are documented in Panel (a) to Panel (c), respectively. As shown, the measured SMDs were all at a level of 10–100 µm, and it can be seen that the SMD of the surrogate 1 was the smallest under all test planes and all test temperatures, which reflects that a higher volume of flash boiling bubbles can indeed enhance the breakup performance of the fuel spray. Furthermore, as seen in Panel (a) at the fuel temperature of 122 °C, the SMDs of surrogate 2 and surrogate 3 were close to one another. Since the bubble fractions for the two surrogates were also close to each other, it is also evident that the in-nozzle bubbles dominate the liquid atomization in the near-nozzle field. On the other hand, as seen in Panel (a) of Fig. 14, with the fuel temperature of 122 °C, from the axial position of y = 12 mm to y = 18 mm, the SMD decreased as the spray droplets moved downstream. The reason is that large droplets are broken up again (secondary breakup) while interacting with the surrounding ambient air, and the collision of droplets was limited so droplet size did not increase in this measurement regime. While in Panel (b) of Fig. 14, when the fuel temperature reached 132 °C, the SMDs for surrogate 2 and surrogate 3 decreased as droplets moved downstream, but the opposite trend was seen for surrogate 1. These phenomena can be explained by spray droplet collision theory [34,35]. When the droplet size is small enough, the collision of droplets will lead to larger droplets. Similar result is seen in Panel (c) of Fig. 14. For surrogate 3, because the fuel droplets were relatively larger, the interacting force between fuel droplet and ambient air was more significant than the influence of droplet collisions. Then, the SMD of the surrogate 3 spray kept decreasing. Meanwhile, when the droplet size was larger, the collision of droplets would not increase the size of the droplet [35]. For surrogate 1, because the droplet diameter was smaller, the mass gained from droplet collisions was more significant than the mass loss from fuel-air interacting. For surrogate 2, the two factors competed with each other, therefore the droplet size would increase first, then decrease later. To further analyze the spatial distribution of spray droplet diameter, we also investigated the SMD of the flash boiling sprays at various 65

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(a)

(d)

Tf =122 Axial position=12mm

Tf =132

Axial position=12mm

Tf =132

Axial position=15mm

Axial position=15mm

Axial position=12mm

(h)

Tf =143

Axial position=15mm

(f)

(c)

Tf =122

Tf =143

(e)

(b)

Tf =122

(g)

Axial position=18mm

Tf =132

Axial position=18mm

(i)

Tf =143

Axial position=18mm

Fig. 15. SMD as a function of radial position for multi-component fuel surrogates under different temperatures. Panels (a)–(c): Fuel temperature of 122 °C, Panels (d)–(f): Fuel temperature of 132 °C, and Panels (g)–(i): Fuel temperature of 143 °C.

Tf =122 Surrogate 1

(g)

(d)

(a)

Tf =132 Surrogate 1

Tf =143 Surrogate 1 (h)

(e)

(b) Tf =132 Surrogate 2

Tf =122 Surrogate 2

(c)

Tf =143 Surrogate 2

(i)

(f)

Tf =122 Surrogate 3

Tf =132 Surrogate 3

Tf =143 Surrogate 3

Fig. 16. Multi-component fuel spray droplet number distribution in different droplet size regions (Axial position is 15 mm). 66

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4. Conclusion

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In this work, we have utilized a two-dimensional transparent nozzle to perform high-speed shadowgraph and PDI measurements for various single component fuels and multi-component fuel surrogates spray characteristics. We investigated the impacts of multi-component fuel compositions on the characteristics of in-nozzle flash boiling bubbles and near-nozzle field spray morphology. Furthermore, we analyzed the impact of single component fuel and multi-component fuel surrogates on the stability of flash boiling sprays, which further illustrated the near-nozzle field breakup mechanisms of multi-component fuel surrogates. Experimental results can also be used by numerical simulations for validation purposes. The following conclusions can be obtained: (1) The addition of high volatility fuels into low volatility fuels would increase the bubble fraction in the injector nozzle, which in turn improves the breakup and atomization efficiency of the flash boiling spray. The level of flash boiling is mainly affected by the bubble fraction in the nozzle, which is closely related to the fuel temperature and the local bubble point pressure of fuel mixture. (2) With similar in-nozzle bubble fraction conditions, the flash boiling sprays with a high proportion of low volatility fuels are more stable owing to the heat and mass transfer mechanisms of fuel mixtures in flash boiling state. (3) The addition of high volatility fuels induces a higher level of flash boiling, and consequently finer spray droplets in the near-nozzle field. (4) The SMDs of the flash boiling sprays during downward movement are co-determined by droplet collisions and droplet-air interaction. The central region of the spray seems to receive more impact of droplet collisions and the outer region of the spray has a stronger interaction with the co-flow air. Acknowledgements This research is sponsored by National Natural Science Foundation of China (NSFC) under grants No. 51376119/E060502. It was carried out at the National Engineering Laboratory for Automotive Electronic Control Technology of Shanghai Jiao Tong University. References [1] Kawano D, Senda J, Wada Y, Fuel Fujimoto H. Design concept for low emission in engine systems 4th Report: effect of spray characteristics of mixed fuel on exhaust concentrations in diesel engine. SAE Int 2003. [2] Myong K-J, Arai M, Tanaka T, Senda J, Fujimoto H. An experimental investigation and numerical analysis of multicomponent fuel spray. Jsme Int Journalserb Fluids Thermal Eng 2004;47(47):200–6. [3] Kawano D, Ishii H, Suzuki H, Goto Y, Odaka M, Senda J. Numerical study on flashboiling spray of multicomponent fuel. Heat Transfer – Asian Res 2006;35(5):369–85. [4] Myong KJ, Suzuki H, Senda J, Fujimoto H. Spray inner structure of evaporating multi-component fuel. Fuel 2008;87(2):202–10. [5] Senda J, Wada Y, Kawano D, Fujimoto H. Improvement of combustion and emissions in diesel engines by means of enhanced mixture formation based on flash boiling of mixed fuel. Int J Engine Res 2008;9(1):15–27. [6] Machado GB, da Silva AH, de Oliveira EJ, Barros JEM, Braga SL, Braga CVM, et al. Methodologies for fuel development using surrogate fuels on spark ignition engines.

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