Experimental and numerical investigation on the jet characteristics of spark ignition direct injection gaseous injector

Applied Energy 105 (2013) 8–16

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Experimental and numerical investigation on the jet characteristics of spark ignition direct injection gaseous injector Iman Chitsaz a,b, Mohammad Hassan Saidi a,⇑, Ali Asghar Mozafari a, Alireza Hajialimohammadi b a b

Center of Excellence in Energy Conversion (CEEC), School of Mechanical Engineering, Sharif University of Technology, P.O. Box 11155-9567, Tehran, Iran CAE Department, IranKhodro Powertrain Co. (IP-CO), Tehran 1398813711, Iran

h i g h l i g h t s " We provide the numerical and experimental work on the transient start of injector. " Numerical simulation at the initial stage is not in exact agreement with the experiment. " Nondimensional tip penetration is located between lines with slopes of 2.625 ± 0.875. " Pressure drop at the initial stage is greater than its steady decrease.

a r t i c l e

i n f o

Article history: Received 12 February 2012 Received in revised form 4 August 2012 Accepted 7 November 2012 Available online 11 January 2013 Keywords: Direct injection Jet structure Gaseous injector Underexpanded jet Image processing

a b s t r a c t Natural gas has widely been used as a fuel in conventional Diesel and spark ignition engines. The better understanding of injector parameters on the jet structure is helpful for the combustion optimization. This paper presents an experimental and numerical study on the jet structure of gaseous fuel injector in spark ignition direct injection engine by Schlieren technique and numerical procedure. Helium was injected through a gaseous injector at the different pressure ratios and nozzle diameters to understand the effects of nozzle geometry and pressure ratio for a dedicated correlation of CNG–SIDI injector. It was found that higher pressure ratio and exit nozzle diameter led to more tip penetration except the initial stages of jet development. Numerical simulation at the initial stage of jet development was not in exact agreement with experimental data due to transient effects of the needle lift within the injector tip and experimental errors while reliable results was observed after 1 ms from the start of injection. It is also notable that tip angle of the jet did not have a specific trend when jet develops. Ó 2012 Elsevier Ltd. All rights reserved.

1. Introduction Direct injection in internal combustion engines enables the realization of stratified charges and seems the most promising way [1–3] of improving fuel economy in spark-ignition engines. It is also notable that among the alternative fuels, methane is considered very promising either because it can work with high compression ratios without experiencing the knock phenomenon or because of its clean combustion [3]. Therefore, direct injection of CNG in SI engine has now become a challenging and innovative technology. Spray characteristics of direct injection is very important factor for providing more stable combustion and lower emissions. Therefore, many researchers have focused on the transient injection of fuel as the start of mixture formation in spark ignition [4] and compression ignition [5] engines. ⇑ Corresponding author. Tel.: +98 21 66165522; fax: +98 21 66000021. E-mail address: [email protected] (M.H. Saidi). 0306-2619/$ - see front matter Ó 2012 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.apenergy.2012.11.023

When the pressure ratio between the outlet and inlet of an orifice or nozzle is sufficiently dropped to cause the pressure at the exit to be higher than the ambient, the generated complex flow is termed an underexpanded jet. Transient underexpanded jet is one of the challenging areas due to its behavior and industrial applications. Early studies of gas injection in internal combustion engine confined to the Diesel engines fueled with natural gas. Carlucci et al. [6] investigated the CNG jet structure by means of shadowgraphy technique for a dual-fuel engine. They analyzed the effect of CNG and Diesel fuel injection pressure, together with the amount of fuel injected during the pilot injection. An intensifier injector system for natural gas fueling of Diesel engines was developed by Hodgins et al. [7] and Gebert et al. [8]. Zhang et al. [9] performed numerical simulation of gas injection of a Diesel engine. Hodgins et al. [10] developed an injector for simultaneous injection of high pressure natural gas and some quantity of liquid diesel. Mezo [11] and Chan [12] investigated stratified charge engine operation using direct injection combustion bomb and employing

I. Chitsaz et al. / Applied Energy 105 (2013) 8–16

9

Nomenclature CI compression ignition CMOS complementary metal oxide semiconductor CNG compressed natural gas CNG–SIDI compressed natural gas spark ignition direct injection d diameter DI direct injection f focal length of concave mirror h thickness K Gladstone–Dale’s constant LIF laser induced florescence M Mach number P pressure RNG renormalization group SI spark ignition SIDI spark ignition direct injection t time U velocity

Schlieren method. They numerically and experimentally focused on the injection and combustion of natural gas and used a modified injector for direct injection of gaseous fuel. White and Milton [13] studied natural gas injection system, based on experimental and numerical analysis of high-pressure injectors to optimize injection timing and injector geometry. Their numerical results were calibrated with the experimental data in order to apply CFD models for optimization. Ouellette and Hill [14] presented a semi empirical model for predicting the underexpanded gas penetration. They experimentally and semi analytically determined the effects of pressure ratio, lift, injection duration and wall constraints on the jet. They validated their model and showed that their predictions had a good agreement with the experimental data. Abraham [15] presented the theoretical analysis of transient gas jets in a quiescent ambient environment, in density ratios of 2, 1, and 0.5. Lahbabi et al. [16] and Johari et al. [17] presented a correlation for the tip penetration of gaseous jets. Abraham et al. [18] investigated on the tip penetration measurement of direct injection of methane and tetradecane through a multi-hole injector in a Diesel engine. Ishii et al. [19] numerically and experimentally investigated the time evolution of circular pulse jets. They showed that unsteady second shocks were realized for all sonic underexpanded jets. Lacerda [20] investigated the start of underexpanded jet. Lacerda studied the effect of pressure ratio as well as the injected gas composition, and also concluded that the shape of the jet was the function of pressure ratio and the kind of gas. Ouellette and Hill [21] numerically studied three dimensional transient jets using k-e turbulent model for grids typical of those used in engine simulations. In their work, they discussed other factors affecting the accuracy of the calculations. Radulescu and Law [22] investigated the initial transient hydrodynamic evolution of underexpanded slit and round jets. They presented a closed-form analytic similarity solution for the temporal evolution of jet while twodimensional numerical simulations were performed to investigate the flow field. The simulations confirmed the similarity laws derived theoretically and there was a good agreement between results obtained from their model, the numerical simulations and the experiment. Recently Chitsaz et al. [23] developed a semi analytical solution for transient start of underexpanded jet using Hankel and numerical Laplace transform. Peterson [24] studied the transient injection of helium and hydrogen that was injected by the prototype multi-hole gaseous injector. Mohamad and Geok [25] examined a spark plug fuel injector which was a combination of fuel injector and spark plug by means of LIF method to understand

Z

tip penetration

Greek symbols line slope b constant d uncertainty c specific heat ratio q density H jet angle

a

Subscripts a ambient n nominal s static t time x in axial direction

tip penetration of their new injector. Nozzle exit reflector on a supersonic jet discharged from a convergent–divergent nozzle was investigated by Kweon et al. [26]. They used a high-quality spark Schlieren optical system to visualize detailed jet structure with and without the reflector. They concluded that in the overexpanded jets, the reflector substantially increases the jet spreading rate and reduces the supersonic length of the jet. Otobe et al. [27] investigated the detailed near-field structures of highly underexpanded sonic free jets using computational fluid dynamics. They presented a correlation for the diameters of the Mach disk that implied the near-field structure of highly underexpanded sonic free jet was a unique function of the pressure ratio, regardless of the nozzle geometry. Computational modeling of the gas injection process in a large-bore was investigated by Li et al. [28]. They examined the effect of the pressure ratio, with ratios ranging from 3.5 to 80, on the velocity and pressure profiles in the near field region. Their results showed good agreement with the results from the method of characteristics applied to two-dimensional jets. In the previous works, experimental and numerical simulation of natural gas injection has been considered in detail for the Diesel engines. Injection pressure of gaseous fuel in Diesel engines is at least twice of those CNG–SIDI engines [29,30]. It is also notable that the pressure ratio and the jet configuration are quite different in these two types of engines. High injection pressure and small nozzle diameter is applied in Diesel engines while pressure ratio is limited and different nozzle diameters have been studied in this work. This paper has focused on the gas injection in CNG–SIDI engine that has not already been addressed in the literature. The present study has two major objectives. The first is to find the jet structure of a new gaseous injector for a SIDI engine at different pressure ratio and nozzle diameters experimentally and the second is the development of numerical effort in jet modeling of SIDI engine. The jet structure was visualized by Schlieren method while numerical simulation was performed by KIVA3V [31]. Helium was injected at pressure ratios ranging from 1.5 to 3.5, resulting in jets spanning from subsonic to underexpanded condition. Tip penetration and the jet angle were also measured to understand the effect of inlet conditions of the gaseous injector. 2. Mathematical model In the present study, simulations were performed by KIVA3V. A second order upwind scheme was used to discretize both momentum and continuity equations. The turbulent flow was modeled by

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I. Chitsaz et al. / Applied Energy 105 (2013) 8–16

Fig. 1. Sketch of experimental setup.

Table 1 Experimental condition. Ambient temperature (K) Ambient pressure (MPa) Injector hole diameter (mm) Injected gas Pressure ratios Injection period (ms) Interval time (ms)

293 0.0872 0.5, 1.0, 1.5, 2.0 He 1.5, 2.0, 2.5, 3.0, 3.5, 4.0 10 0.2

a modified renormalization group (RNG) k-e model [32]. Gas injection in KIVA3V was modeled by Ra et al. [33] procedure. 3. Experimental apparatus The experimental apparatus includes fuel injection system, jet imaging system and sensors which are shown in Fig. 1. An injector controller sends a signal for injection and monitors pressure and temperature of injection line. The temperature and pressure of the injection medium in the experiment are set to 0.87 bar and 293 K. The fuel jet imaging system is comprised of concave mirrors (f = 2610 mm and f = 2570 mm), halogen lamp, CMOS camera, and a knife edge. High speed motionblitz cub3m3 CMOS camera, with a maximum resolution of 512  512 at 2500 fps and a AF Nikon 70– 300 mm lens were used. Pressure sensor was jumo dtrans p30 and calibrated by dpi pressure calibrator of druck company. Thermometer was ds-k-tf of BOSCH company and mounted in the line of gas injection. The gas flow meter of testo6441 was applied to measure the flow rate. The injection nozzle had the diameter of 0.5–2 mm and was capable of supplying injection pressures up to 3 MPa. The helium was supplied from a standard, 20 MPa tank. For consistency and repeatability of testing, the injector’s supply pressure was limited to 2 MPa with an in-line regulator. Experimental conditions are summarized in Table 1. 4. Image processing Jet visualization was used to investigate the flow structure of gaseous jet produced by the prototype injectors. Jet patterns were

imaged both vertically and end on or directly towards the tip of the injector. The jet tip penetration was measured using gradient of pixel values. Edges of the jet were specified by looking for local maxima of the gradient of pixel values. The vertical penetration was then defined as the location where the gradient of intensity fell below a certain value, starting from the bottom of an image. This procedure also can be done using edge detection in MATLAB. Fig. 2 further illustrates this technique. In each image, injector position was specified, and then tip penetration angle with its magnitude was calculated. 5. Results and discussion Numerical simulations were performed to evaluate the capability of reproducing transient jet characteristics. Helium was injected at a pressure ratio of 1.5–3.5 into air through different nozzle diameters. Injection medium was filled with air at a pressure of 0.872 bar. The refractive index of helium is 1.000035 while refractive index of air and methane are 1.00292 and 1.00044, respectively. This minor difference in refractive index has a great impact on the visualization of small gas density variations in Schlieren method. So Helium has been selected for this experiment to enhance the quality of the captured images. To avoid grid-size dependency; four mesh sizes with the same conditions were used to check mesh dependency of the problem. Helium mass fraction of axisymmetric line of jet for different mesh sizes are given in Fig. 3. As it is shown, helium mass fraction along axis does not vary after refinement of 16,500 meshes. Fig. 4 shows the helium density contours of the present computational results with the experimental ones. The Schlieren visualization is shown in the lower half, while the present computational results are given as predicted isodensity contours in the upper half. The pressure ratio and diameters are 1.31 and 1 mm respectively and time set to 4 ms after start of injection. As it is shown in this figure, tip penetration is predicted quite accurate in the numerical simulation. Jet development in the radial direction is also shown in this figure. Jet width has a local maximum in numerical results which is shown in the figure and has some deviation with that of experiment. As is shown in this figure, due to turbulent and molecular mixing, the flow pattern that exits from nozzle in a supersonic regime tends to collapse towards the axis immediately, leading to

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I. Chitsaz et al. / Applied Energy 105 (2013) 8–16

Fig. 2. Description of penetration measurement by edge detection.

120

1 0.9

100

0.7 0.6

Tip Penetration (mm)

Helium Mass Fraction

0.8

17000 cell 16500 cell 15500 cell 15000 cell

0.5 0.4 0.3 0.2 0.1 0

0

0.02

0.04

0.06

0.08

0.1

0.12

0.14

0.16

80

P/Pe=3.07

60

40 Exp. Numerical

P/Pe=1.31

20

Axisymmetric Line Fig. 3. Helium mass fraction along axisymmetric of jet at pressure ratio 1.31 for different mesh size.

0

0

1

2

3

4

5

6

Time (ms) Fig. 5. Tip penetration comparisons of experimental and numerical data at d = 1 mm.

Fig. 4. Helium density of simulated and experimental results in pressure ratio 1.31 and d = 1 mm at 4 ms.

the typical bell-like shape of the jet. The needle-tip opening law plays a very important role in the jet shape definition. In fact, mass-flow rate fluctuations can arise if the needle opening is too fast, leading to problems of jet stability and control of local air fuel ratio in the combustion chamber for stratified engine operations. Tip penetration in the numerical simulation was defined as the distance from the injector’s exit to the fuel mass fraction within the gaseous plume had decayed to 3%. Defining the jet tip penetration using 5% and 1% as mass fraction criteria does not exhibit important differences in the penetration curves [34]. Fig. 5 shows the penetration rate of the simulated jet and the corresponding experimental results at d = 1 mm. As it is shown in this figure, the initial stage of jet development is not captured well in the numerical sim-

ulation but after a 1 ms, the numerical perdition comes close to experimental data. This trend mismatch at the initial stage is created due to time resolution of images that was set to 0.2 ms or transient effects of the needle lift within the injector tip and disappear during the jet development. Tip penetration measurement of two different nozzle diameters at pressure ratios ranging from 1.5 to 3.5 are shown in Fig. 6. As shown in this figure higher pressure ratio leads to more tip penetration when the jet propagates enough. The jet develops throughout its duration but more rapidly at the beginning of the process. This leads to more deviation between experimental and numerical results at the initial stage of jet development due to time resolution for the start of jet. Experimental errors have more influence on the tip penetration measurement at the early times of injection and become infinitesimal compare to jet penetration when jet propagates enough. This indicates that measurement greater than 1 ms is more reliable than those of early steps of injection. This figure indicates that tip penetration is not highly affected by pressure ratio for smaller diameters. The smaller diameter will reduce the time available for the fuel–air mixing, easily to bring an insufficient fuel–air mixing and decrease the mixing quality between air and fuel. In addition, lower pressure ratio will also decrease the jet penetration distance for the late part of injected fuel after intake valve

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I. Chitsaz et al. / Applied Energy 105 (2013) 8–16

0.1

0.12

0.09 0.1

Tip Penetraton (m)

0.08

Tip Pentraton

0.08

0.06

0.04 P/P0=1.5 P/P0=2.0 P/P0=3.0 P/P0=3.5

0.02

0

0

0.5

1

1.5

2

2.5

3

3.5

0.07 0.06

P/P0=2.5 P/P0=3.5 P/P0=2.0 P/P0=1.5

0.05 0.04 0.03 0.02 0.01 0

4

0

1

2

3

Time (ms)

4

5

6

Time (ms)

(a) d=1mm

(b) d=0.5 mm

Fig. 6. Tip penetrations of several pressure ratios at two different diameters.

100

100

d=0.5 d=1.0 d=2.0

80

P/P0=3.5 P/P0=2.5 P/P0=1.5

80

60

40

(deg)

(deg)

60

40

20 20

0

0

-20

-40

0

0.5

1

1.5

2

2.5

3

3.5

4

-20

0

0.5

1

1.5

Time (ms)

2

2.5

3

3.5

4

Time (ms) (b) d=0.5 mm

(a) P/P0=1.5 Fig. 7. Tip penetration angle at different diameters and pressure ratios.

closing, making more fuel to be concentrated into the range near the injector exit, and this will bring a unstable combustion due to the decrease of rich mixture around the spark plug. This is very important factor for spray guided mixture formation that injector have the fundamental effect on mixture distribution. Fig. 7 shows the transient jet angle measurements associated with the iterative method of penetration measurement for several pressure ratios and diameters. Jet angle, h, is defined as the angle of tip penetration into injector axis. This angle is higher at early times due to the development of the head vortex and then it drops to a relatively steady value. The transient jet angle measurements show a significant amount of scatter and inconsistencies between individual runs at the same condition. For a majority of cases it is hard to identify a single, steady-state value of jet angle but it comes to zero when jet develops in the time. It is also notable that pressure

ratio and also diameter do not significantly affect on the jet angle tip. For most cases, turbulence behavior of jet does not give the predicted angle for specific pressure but all cases lie in the ±5° to the axis of injection. When high pressure gas is injected into the lower pressure medium, gas accelerates and its static pressure falls. Fig. 8 shows the gauge pressure of injection line. As is shown in this figure, initial pressure drop is greater than its steady decrease due to the initial acceleration of jet. When the nozzle exit diameter increases, more flow is passed through the injector. This results in more pressure drop for bigger inlet nozzle diameters. For a compressible flow, the static pressure can be presented as a function of its total pressure and its exit Mach number according to Eq. (1). Helium with a specific heat ratio of 1.66 has the critical pressure ratio of

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I. Chitsaz et al. / Applied Energy 105 (2013) 8–16

1.5

d=0.5 mm

3.2

d=1

d=2mm

2.8

1.3

2.6

Pressure (bar)

Presure (bar)

d=1.5

1.4

3

2.4 2.2 2 1.8

1.2

1.1

1

0.9

1.6 0

1

2

3

4

5

6

7

8

9

0.8

0

1

2

3

4

5

Time (ms)

Time (ms)

(a) P/P0=3.5

(b) P/P0=1.5

6

7

Fig. 8. Effect of pressure ratio and nozzle exit diameter on the in-line injection pressure.

120

Tip pentraton (mm)

100

80

60

40 d=0.5mm d=1mm d=2mm

20

d=1.5mm

0

0

0.5

1

1.5

2

2.5

3

3.5

4

Time (ms) Fig. 9. Effect of nozzle exit diameter on the tip penetration at P/P0 = 1.5.

Fig. 10. Nondimensional penetration rate of turbulent Helium jets for different nozzle diameters and pressure.

2.05. At higher pressure ratios, the exit Mach number remains equal to one, but the exit static pressure is greater than the ambient or cylinder pressure. Thus a substantial gas expansion from the exit pressure to the cylinder pressure is required towards downstream of the nozzle. This type of nozzle outflow produces an underexpanded jet. For methane injection, the critical pressure ratio equals to 1.85 and underexpanded condition happens earlier.



ps ¼ pa 1 þ

c1 2

M

2

1c c ð1Þ

Fig. 9 shows the tip penetration for the different nozzle diameters. As it is shown in this figure, larger nozzle diameters lead to more tip penetration at a constant pressure ratio. It is also notable that variation of tip measurement with diameters is not linear

while the linear increase can be seen for increment of the pressure ratio. At the initial stage of jet development, tip penetration for different nozzle diameters does not follow its increasing trend for higher inlet nozzle diameters due to those mentioned reasons for the different pressure ratios. Slight changes of tip penetration is observed from 1.5 mm to 2 mm of inlet nozzle diameter and indicates that 1.5 mm would be optimum for spray guide mixture formation in CNG–SIDI engines. Previous investigators [14–18] have shown that the proper characteristic time and length scales for jet measurement are in the form of Eq. (2).

Z dn ðqn =qa Þ1=2

¼C

p1=4 4

tU n dn ðqn =qa Þ1=2

!1=2 ð2Þ

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I. Chitsaz et al. / Applied Energy 105 (2013) 8–16

in which dn, Un, and qn are the nozzle diameter, exit velocity, and exit flow density, while qa is the chamber density. As shown in Eq. (2), tip penetration is proportional to the fourth root of the injected gas density. So tip penetration of CNG can be calculated by product of fourth root of density ratio ([qCNG/qHe]1/4) in the helium tip penetration. By defining non-dimensional time and tip penetration in Eq. (3), Eq. (2) can be rewritten as Eq. (4).

zn ¼

zt dn ðqn =qa Þ1=2

;

tn ¼

!1=2

tU n

ð3Þ

dn ðqn =qa Þ1=2

zn ¼ at 0:5 n

ð4Þ

Fig. 10 shows the nondimensional tip penetration versus nondimentional time data for the current injector. All data is located between lines with slopes of 2.625 ± 0.875. This slope value is obviously different from the range of previously reported values between 2.8 and 3.1 [13–18]. This trend indicates that tip penetration in CNG–SIDI injector has a lower tip penetration than those of used for Diesel engines. Due to low cetane number of methane, natural gas does not spontaneously ignite under typical CI compression ratios (and corresponding temperatures) like Diesel, but needs a source of controlled ignition. Thus it should be injected

at the pressure range of pilot fuel injection. This results in a shortened ignition delay of the pilot fuel [35] because of faster mixing between the pilot fuel and air during the ignition delay period while Natural-gas fueled SI engines can operate at lower pressure ratios of injection due to the stratified mode of mixture distribution. In these engines rich mixture near spark is needed for the start of combustion and then flame propagates to the leaner portions. Gaseous injector in CNG–SIDI engines should guide the gas to the spark plug at the proper time and it can be seen by lower pressure ratios and bigger nozzle diameters in comparison with Diesel engines. It is also notable that tip penetration normalization is only valid for distances greater than about 20 nozzle diameters. This limitation causes some of normalized data lie out of lines. 6. Error analysis Due to the experimental errors in the measurement of tip penetration, an error analysis should be performed to understand the uncertainty of reported results. For a parameter of interest f(xi), which is a function of several measured parameters xi, the overall uncertainty is represented by Eq. (5). d is an uncertainty of experimental apparatus or experimental procedure [36].

0.4

0.7

0.35

0.6 0.5

Relative Error

Relative Error

0.3 0.25 0.2 0.15

0.1

0.05 0

1

2

3

4

5

6

0

7

0

0.35

0.6

0.3

0.5 0.4 0.3

5

0.15

0.1

0.05

2

2.5

3

3.5

7

0.2

0.1

1.5

6

0.25

0.2

1

4

d=1

0.7

0.5

3

d=0.5 0.4

0

2

Time (ms)

0.8

0

1

Time (ms)

Relative Error

Relative Error

0.3 0.2

0.1

0

0.4

0

0

0.5

1

1.5

2

2.5

3

Time (ms)

Time (ms)

d=1.5

d=2.0

Fig. 11. Estimated error of different nozzle diameters at pressure ratio = 3.0.

3.5

4

4.5

5

I. Chitsaz et al. / Applied Energy 105 (2013) 8–16

D¼

(  2 N X @f @xi

i

)0:5 d2xi

ð5Þ

As it was mentioned, tip penetration was represented by Eq. (2). Therefore uncertainty of tip measurement can be represented by Eq. (6).

(

D¼

@f @dn

2

d2dn



@f þ @ qn

2

 2 )0:5  2 @f @f 2 þ dt þ d2pn @t @pn

2

dqn

ð6Þ

All dimensions had a precision of 0.1 mm. So uncertainty of dimensions was set to 0.05 mm while pressure sensor had 0.001 bar precision. In the present experiment, time interval of captured images was 0.2 ms and uncertainty of time equals to 0.1 ms. uncertainty of density is due to Schlieren experimental setup and is determined by Eq. (7)[37].

qx ¼

@q dx dx2 ¼ ) dq ¼ @x f2 hK f2 hK

ð7Þ

where q is the gas density, h is the thickness of inhomogeneous medium under test in the ray propagation direction and K is the Gladstone–Dale’s constant. The Gladstone–Dale’s constant is a function of both the wavelength of the light source and the physical properties of the gas. In the present study, its value was taken as 1.96  104 m3/kg, for the helium temperature of 293 K [38]. Derivatives in Eq. (6) is calculated by tip penetration formulation and represented in Eq. (8). 1

@f @dn

1=2 ¼ 12 dn bt 1=2 ; @@fq ¼ 14 q1 ¼ 14 n bt

@f @pn

@ qn @U n 1=2 1=2 @U n 1=2 ¼ 12 U 1 þ 14 q1 ¼ 12 U 1 þ 14 n @p bt n @p bt n @p bt

n

n

n

RT n pn

bt 1=2 ; @f ¼ 12 t 1 bt 1=2 @t n

1 RT n

bt 1=2 ð8Þ

p 1=4

where b in Eq. (8) equals to C 4 ðdn ðqn =qa Þ1=2 U n Þ0:5 . Fig. 11 shows the estimated errors of different nozzle diameters when jet propagates in time. As shown in this figure, at the early steps of jet development relative error is as much as 75% and decreases to lower than 10% after 1 ms. This error also is reduced continuously in a slower rate after 1 ms from start of injection. In most cases, relative error is decreased to 5% after 3 ms from start of injection. Larger error at the beginning of injection is due to limited interval time of recorded images. The uncertainty arising from time increases the level of error before 1 ms while uncertainty from positioning of Schlieren setup is more important after 1 ms. Maximum relative error in this experiment is lower than 5% after 2 ms from start of injection. 7. Conclusions The present study provides the numerical and experimental work to investigate the transient start of new injection system for the direct injection of gaseous fuel in SIDI engine. Different nozzle diameter and pressure ratios were investigated to understand the effect of inlet condition of gaseous injector and dedicated correlation of CNG–SIDI injector. Both numerical effort and experiment have been done to understand the behavior of this type of injector. Numerical simulation at the initial stage of jet development was not in exact agreement with experimental data due to transient effects of the needle lift within the injector tip and experimental errors while reliable results was observed after 1 ms from the start of injection. It is also notable that relative error was as much as 75% at the initial stage of jet development while it was reduced to below than 5% after 2 ms from start of injection. Larger error at the beginning of injection is due to limited interval time of recorded images.

15

It is concluded that at the higher pressure ratios and nozzle exit diameters, more tip penetration is observed when the jet propagates enough. It is also notable that the variation of tip measurement with diameters is not linear while linear increase can be seen for increment of pressure ratio. It was shown that pressure drop at the initial stage is greater than its steady decrease due to the initial acceleration of jet. It is also evident that for the bigger nozzle diameters, pressure drop is higher due to more passing momentum. Jet angle of tip was measured respect to the axis of injection and no specific trend was observed for different pressure ratios and nozzle exit diameters. The transient jet angle measurements showed a significant amount of scatter and inconsistencies between individual runs of the same condition. For a majority of cases it was hard to identify a single, steady-state value of jet angle but it came to zero when jet was fairly developed. It was shown that all nondimensional data was located between lines with slopes of 2.625 ± 0.875. This slope value is obviously different from the range of previously reported values between 2.8 and 3.1. This trend indicates that tip penetration in CNG–SIDI injector has a lower tip penetration than those of used for Diesel engine.

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