Electrostatic Atomization and Combustion of Electrically

Insulating Liquids

John Shrimpton1 University of Southampton, UK

and

Farzad Mashayek2

University of Illinois at Chicago, USA This review covers the literature and patents relevant to a specific type of liquid and

a specific method of atomization. The liquids are dielectrics; poor electrical conductors, typically vegetable oils such as corn, soy and rapeseed, or petroleum products, such as petrol/gas, aviation fuel and Diesel oils. The liquids need not be ‘doped’ to enhance their electrical conductivity. The ‘charge injection’ atomizer injects electric charge into the poorly conducting liquid and the liquid atomization, spray dispersion and combustion are significantly influenced by the presence of the injected electric charge. The review surveys studies of the primary atomization process, spray characterization, and effect on combustion. The review concludes that whilst some fundamental understanding still requires more research, sufficient knowledge exists to design and operate practical devices.

I. Introduction and Scope

The creation and atomization of electrically charged insulating liquid jets and the dynamics of the charged

sprays so produced are not subjects that have been widely reported in the literature. Such a technique, if widely available, could well be valuable in a number of applications, for instance in molten plastic production and a range of fuel spray combustion systems. The benefits of electrically charged fluid mechanics are well known and are used successfully in a range of industrial applications, such as production of nano-particles1, to spray coating applications and flue gas treatments2 and include; 1) Low drop concentration within the plume. 2) Lack of drop agglomeration. 3) Controllable and constrained particle size distribution. 4) Controllable spray plume shape and deposition.

These benefits, unique to electrically charged sprays, are also attractive for combustion applications, where a few large drops and regions of high drop concentration induce non-uniform combustion kinetics within a burner. This in turn leads to non-optimum combustion and production of unwanted products. Historically it has been problematic to extract charge from the charging electrode, a metallic conductor, into an electrically insulating liquid using an electrostatic atomization method. The traditional way to circumvent this problem has been to add small amounts of an additive that would reduce the electrical resistivity, from a high value of ~1010 Ωm, typical of an insulating liquid, to a more useful range, typically ~107 Ωm. The doped insulating liquid, from an electrical viewpoint now ‘semi-conducting’, permits the liquid to be atomized using standard contact charging techniques. For further information on this, and the induction charging of drops, the reader is referred to the work of Law3. For an explanation of the role of electrical conductivity in electrostatic atomization see Ganan-Calvo et al.4. A range of hydrocarbon oils have been sprayed by doping the fuel5-7 and although the flow rates were very small, limited by the atomization method, combustion was readily achieved. Research has also suggested that the presence of electric charge on the drops that form the spray plume enhances the overall evaporation rate8 and also may reduce the amount of soot produced9 due to the increase of inter-particle

1 Senior Lecturer, Energy Technology Research Group, School of Engineering Sciences, Southampton, S017

1BJ. 2 Professor, Department of Mechanical and Industrial Engineering, 842 W Taylor Street, Chicago, IL 60607,

Associate Fellow of AIAA.

1 American Institute of Aeronautics and Astronautics

47th AIAA Aerospace Sciences Meeting Including The New Horizons Forum and Aerospace Exposition5 - 8 January 2009, Orlando, Florida

AIAA 2009-667

Copyright © 2009 by the authors. Published by the American Institute of Aeronautics and Astronautics, Inc., with permission.

separation. The use of corona is known to reduce NOx emissions of natural gas flames10, and extraction of soot by applying an external field across a flame has been demonstrated by Lawton & Weinburg11.

The charge injection method has evolved from the pioneering work of Kim & Turnbull12 who sprayed Freon 113 using a needle with a chemically etched tip radius less than 1μm, which was placed in a glass capillary. However only extremely small flow rates and currents of 10-3 ml/s and 10-9 A respectively, could be attained. In addition the atomization performance quickly degraded because the needle tip became blunt due to the high current flux emitted. This technique should not be confused with the addition of inserts for spraying semi-conducting liquids, as proposed by Cloupeau13 where the intention was to assist the formation of a small liquid cone and not to supply electrical charge. For charge injection devices a central electrode is maintained at a negative potential, typically 10kV magnitude, and is the source of electric charge in the liquid. The charge is ‘injected’ due to the strong and highly non-linear electric field present at the emitter tip but the mechanism is as yet not fully understood. The low flow rate limitation12 was improved by Kelly14-16 who made the atomizer orifice contraction an additional earthed electrode and denoted his patented design a 'Spray Triode'. Instead of using a single charge injection site12 a eutectic material17 was used and it was claimed that this acted as a vast array of injection points, with the typical tip radius less than 1μm. Using this approach the atomization performance and more importantly the flow rate of the ‘Spray Triode’ were not limited by the charge supply rate because the current flux injected, per tip, was very small, and therefore the electrode did not degrade. Kelly16 proposed that the 'Spray Triode' design operated essentially as a massively multiplexed version of the single point concept pioneered by Kim and Turnbull and therefore that an electron/ion mechanism was responsible for the charge injection.

The content of the review is taken primarily from Shrimpton18, Rigit19, Shrimpton & Yule20-21, Rigit & Shrimpton22. This specific set of references is used to give a concise overview of the basic principles of operation, and is used for the simple reason that complete data is available to the authors. Romat23-25 has developed a similar, higher pressure design. Also of relevance is a set of patents filed by Kelly whilst employed at Exxon14-15,26-29 and two journal publications16,30 relating to his ‘Spray Triode’ device. To conclude the review more specific technology advancements published in the literature and in patent disclosures are reviewed.

Electrostatics is a term used to describe the physics of charge in motion and at rest in the absence of significant magnetic field effects and can be used to refer to any type of phase where this is the case31. Electrohydrodynamics (EHD) is a more specialized term that is generally used to refer to the role of electrostatics in liquid media. A good introduction to electrostatics can be found in texts such as Crowley31 and Chang et al.10 with a more EHD orientated approach taken in Melcher32 and, specific to dielectrics, Castellanos33.

II. Experimental Details

A generic atomizer, Fig. 1, was fabricated in three sections and the conducting metal parts were fabricated from brass for ease of machining and to eliminate corrosion problems. The nozzle sections, containing the version 1 and 2 profiles, Fig. 2, were threaded to allow removal and replacement onto the body of the atomizer, and these internal parts are shown schematically in Fig. 3. The liquid flow inlet was electrically isolated from the atomizer part on the left of the schematic that contained the high voltage electrode mount via a PTFE spacer and nylon bolts. A hole was drilled in the centre of the PTFE spacer of diameter slightly smaller than that of the steel needle electrode such that the needle was movable coaxially in both directions by a micrometer. The fit was such that the liquid seal was maintained up to the maximum liquid supply pressure of 0.6MPa. The needle was mounted on the micrometer via an insulating section of PTFE so that the electrode gap between the needle tip and the inlet to the orifice channel, Li was adjustable. This was done in order to quantify the response of the spray and leakage currents, IS and IL (defined below) to the varying electrode geometry. The charge emitter potential magnitude was varied in the range 0-25kV, after which the likelihood of breakdown along the PTFE surfaces insulating the electrode from the earthed atomizer body increased. A distinct advantage of the charge injection atomizer design is that the high voltage electrode is immersed in the electrically insulating liquid and that it is also completely encased within the atomizer body, which is earthed. This, for combustion applications has the advantage that the possibility of a corona discharge from the high voltage electrode to the ionized combustion environment is precluded. For non-combustion applications, the fact that the high voltage electrode is deeply imbedded within the atomizer body has obvious safety advantages.

Referring to Fig. 2, the geometrical variables are d, the orifice diameter, L0, the length of the orifice channel and Li the distance between the back face of the atomizer orifice channel and the tip of the emitting electrode. This contrasts with the much more compact ‘Spray Triode’ of Kelly34 shown in Fig. 9 of that paper. The atomizers use a very simple charge injection electrode, a stainless steel needle, and this was mounted on a micrometer to allow variation in the electrode gap Li. The electrode tip varied from a radius of 60μm ± 5μm down to 20μm ± 5μm, depending on the needle used.

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In the charge injection process, a negative DC voltage V is applied between a sharp charge-emitter (the needle) at a point (the tip) of known radius rp and a charge-receiver plate, a plane translated a distance Li from the point, resulting in an injection current IT. Figure 2 shows the key differences between the various internal geometries of the nozzle versions. The version 1 nozzle design was based on the work of Jido35-36 and was reasonably effective providing the orifice diameter was large (~500μm). Smaller orifice diameters presented problems, as discussed below, and a ‘version 2’ geometry similar to Kelly16 was employed. This proved more successful and the ‘version 3’ design provided improved co-axial alignment between electrode tip and orifice, and permitted successful atomizer operation down to ~100μm. Classical point-plane atomizers of d<100μm are extremely difficult to operate successfully, though other designs are capable of this, and will be discussed later.

A typical experimental layout and electrical connections for producing a continuous charged spray of kerosene, white spirit and Diesel oil are shown in Fig. 4. A pressurized vessel was used as a reservoir for the liquid and for inducing liquid flow in the system. The liquid was filtered using sintered metal filters of nominally 7 μm porosity and a standard needle valve was used to maintain the required steady flow, monitored with a rotameter of range 0.10 ≤Q≤ 0.35 ml/s, at worst ±6.0% and 0.15 ≤Q≤ 2.0 ml/s, at worst ±9.0%. Nylon pipe work and push-in fittings connections ensured effective isolation of the nozzle from ground. Two high voltage supplies were used during different phases of the research. A switchable polarity Brandenburg high voltage DC supply (V=±0-30kV) and a Spellman Model SL300PN high voltage power supplied negative voltage up to –24 kV ±1.0%. A grounded sink captured the charged spray and this measured the “spray” current IS, via a Keithley Model 6514 electrometer with an accuracy of ±0.1%. The error arising from not capturing all the spray current was assumed to be negligible and the fluctuations were inherently damped by applying a 10 point moving average to the incoming measurement for the displayed current. The leakage current IL is defined as the current that passes through the liquid, inside the nozzle, to the grounded inner surfaces of the nozzle body and measured with an accuracy of ±7.4% via a Wavetek Model 23XT digital multimeter. The upstream conduction of electric charge was assumed to be negligible.

Regarding leakage current measurement there was a necessary compromise between accuracy and robustness of the instrument since when ‘probing’ the limits of operation of the atomization device large current transients were occasionally observed, leading to tripping of the power supply. This compromise was possible due to the fact that the leakage current is of less importance than the spray current in determining atomizer characteristics.

Several test liquids, physical properties of which are listed in table 1, were used during the experimental program18,19 with a range of viscosities and purities. A common UK domestic heating kerosene fuel marketed by Shell as 'Homeglow 28' was used for the majority of the research. This has chemical, physical and electrical properties similar to saturated hydrocarbons with a molecular weight around 140 (~C10). Standard UK diesel oil (~C12) was also used to provide information on the effects of physical property variation. Tests were also carried out using commercially available ‘white spirit’ which is a low grade UK cleaning solvent with a mixture of aliphatic and aromatic hydrocarbons, the mass fraction of the latter being around 18%.

It was noted during preliminary tests that if a liquid was recycled several times it became easier to charge, which gave an indication that molecular degradation occurred and that the charge injection mechanism is electrochemical in nature. This enforced the experimental restriction that all liquids should be sprayed only once. The resistivities of the kerosene and diesel oil were found, by measuring currents between planar electrodes, to be relatively invariant to the value of the DC potential up to 20kV, as shown in Fig. 5. There was a non-linear response at the lower applied potentials for white spirit and this is thought to be due to the diversity of molecular structures that comprise this liquid, and differences in their respective ion mobilities. This was largely irrelevant since the regime of atomizer operation of most interest generally used potentials of ⏐V⏐≥6kV where the resistivity profile is relatively flat. The relatively lower resistivity of the diesel oil is thought to be due to the presence of traces of detergents and other additives, however all three liquids could be reasonably classified as ‘insulators’ based on these values of resistivity.

III. Breakdown Limits and Typical Current-Voltage Response

The voltage-current response for charge injection atomizers was investigated by first defining a given atomizer geometry and flow through the device. The high voltage was progressively increased and the spray and the leakage current response was recorded, and the total current is assumed to be the sum of these. It was found that two mechanisms for limiting the maximum current, defined by different types of electrical breakdown were present. These, and a typical voltage-current relationship for each, are now discussed. A. Sub-Critical Breakdown

Figure 6 shows the total emitted current, IT, plotted against the potential applied to the charge injection electrode, V, for kerosene, white spirit and Diesel oil for selected flow rates in a ‘sub-critical’ regime, for a

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‘version 1’ nozzle design (Fig. 2). The terminations of the curves are highlighted by circled data points and the loci of these curve terminations for each liquid over the voltage range are represented by dashed lines. Each of the curve terminations is the result of a high voltage breakdown event. Two issues are discussed in this section (1) the voltage current response as a function of liquid properties and (2) the form of the ‘sub-critical’ breakdown event.

It is observed that the total emitted current is a function of applied potential, but not of the bulk velocity. It is also observed that the current-voltage response is a function of the liquid properties. While the charge injection process nominally is only a function of Li and the electrode tip radius the charge injected into the fluid in the small volume near the high voltage electrode tip, it must also be extracted from this region at the same rate. Here the ion mobility has an important role to play, and as such controls the overall rate of charge injected. This is reproduced in Fig. 6, with higher viscosity liquids carrying smaller total currents for a given voltage, independent of applied bulk flow rate.

Evidence of the effect that differing ion mobilities have can be obtained by examining the spray current plotted against applied voltage as given in Fig. 7, for the data-set of Fig. 6 for white spirit and Diesel oil. The kerosene case has been omitted on the grounds of clarity and again the dashed circles and lines represent the maximum spray currents and their loci respectively as in Fig. 6. This implies that electrostatic atomization is inherently more efficient, in terms of the fraction of injected current that leaves with the spray, for more viscous fuel oils, and that the technique could have application in the atomization of the heavier fuel oil fractions. Conversely this, together with the fact that the total injected current is larger for less viscous fuel oils implies that there should be a fundamental limit regarding how low a viscosity can be and be successfully sprayed by the charge injection technique.

Turning now to the breakdown mechanism, Fig. 6 shows the maximum current at breakdown is affected by the flow field with higher bulk flow velocities permitting greater current injection prior to electrical breakdown. This charge limiting mechanism is a form of electrical breakdown and is typical of a high voltage insulation breakdown event and occurs when the electric field at a point equals the electrical breakdown strength of the hydrocarbon liquid. Since the electric field passing through the dielectric fluid is dependent on the emitting electrode, the internal geometry and the space charge distribution, which is strongly coupled to the flow field, the location of the breakdown location is difficult to pinpoint in both space and time.

It was observed however that the electrical breakdown event may be approached by either maintaining a given voltage while reducing the liquid flow rate, or by maintaining a liquid flow rate and increasing the voltage applied to the cathode. Both of these operations have exactly the same effect, that is to increase the space charge concentration and this must lead to an increase in the electrical field magnitude inside the nozzle. At the critical point, the space charge concentration is such that the field exceeds the breakdown strength of the insulating fluid and precipitates an electrical short circuit inside the nozzle. This produces a surge in leakage current, the tripping of the power supply, and the effective cessation of the atomization process.

The breakdown event for the sub-critical flow rate charge injection regime is a ‘catastrophic breakdown’, and is due to complete electrical failure of the insulating liquid. For liquids at rest electrical breakdown is known to occur at smaller electric field magnitudes, and therefore the voltage applied, for less dense hydrocarbons37. As shown in Fig. 6 an increase in flow rate delays the onset of catastrophic breakdown to higher applied potential, and this trend is universal for all test liquids. This is despite an increase in the total current at higher voltages which must increase the space charge averaged over the volume of liquid in the atomizer.

The precise cause for the ‘catastrophic breakdown’ for the sub-critical atomizer operation is unknown. Currently it is thought that since the liquids are pressurized to only few bar gauge pressure, and significantly below their critical pressure, vapor cavities form and at elevated voltages streamers form and reach the earthed nozzle body. It is suggested, but not demonstrated, that larger bulk flow rates inhibit the formation of streamers and delay the onset of ‘catastrophic breakdown’.

Investigation of the sub-critical regime is instructive since it has suggested the existence of a maximum ionic mobility and hence a lower limit for liquid viscosity. However it is not of useful engineering significance in terms of atomization performance due to the unstable nature of the EHD physics and therefore atomizer operation, the low maximum injected spray currents possible and the relatively poor atomization of the liquid. The remainder of this review concentrates on the super-critical regime, which is not subject to these problems. B. Super-Critical Breakdown

As noted above, in the sub-critical regime, breakdown occurred before achieving a maximum spray current on a continuous IS-V curve. Experiments were conducted at higher flow rates using the same version 1 atomizer geometry, Li/d=6.6, L0/d=2, d=500μm, as used above, and a new breakdown mechanism was encountered. Here the catastrophic type of breakdown described in the previous section does not occur and a maximum (optimum) spray current is always achievable for any flow rate and liquid. Figure 8a shows the spray current response where in this case kerosene, white spirit and diesel oil were sprayed at a flow rate of Q=2ml/s for the three

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liquids. The least viscous oil, white spirit, exhibits the lowest maximum spray currents and for this liquid the maximum spray current changes more gradually with applied voltage, in contrast to the more extreme responses of kerosene and diesel oil. This is attributed to the diversity of aromatic and aliphatic molecular structures, and commensurately wide range of ion mobilities, present in the white spirit. Figure 8b shows the leakage current responses for the same cases Fig. 8a and in contrast to the sub-critical regime these are relatively unaffected by changes in liquid viscosity.

Before reaching the optimum operating condition for the atomizer (the maximum spray current) both the spray and leakage currents rise steadily. After this condition the spray current is significantly reduced and the majority of the current flows to the nozzle body and to earth, although the spray current stays at a low, non-zero and near-constant value. It has been experimentally determined that for operation in this regime a stable and repeatable voltage-current-flow rate response can be obtained for all three liquids, i.e. the voltage may be increased above, then reduced below the breakdown value repeatedly and the spray responds with no observable hysteresis effect. Unlike for the sub-critical regime, there is no change in the total current, IT, drawn from the power supply after the spray current maximum occurs. This means that there is no disturbance to the charge injection mechanism near the needle cathode when the ‘super critical’ breakdown occurs. This absence of feedback between breakdown and the charge injection itself, has led the authors to refer to this regime as super-critical, as opposed to sub-critical, operation at lower flow rates. It can be inferred that the electrical and flow conditions local to the charge injection point are unchanged and the breakdown event must occur far from the charge injection site. Since within the confines of the atomizer the flow and electrical parameters are highly coupled, this suggests that the breakdown event does not occur inside the atomizer. The character of the breakdown event is completely different to the ‘catastrophic breakdown’ of the sub-critical regime and it is termed ‘partial breakdown’ since from Fig. 8a there is some residual charge in the spray. This means that in contrast to sub-critical electrohydrodynamics, operation in the super-critical regime is inherently robust since the charge injection mechanism is completely decoupled during breakdown mechanism. Corona discharge is present in air around the liquid jet during the partial breakdown event and this was detected as static recorded in the radio frequencies. In addition a faint purple-blue glow was observed under reduced lighting conditions around the liquid jet as it emerged from the orifice. Figure 9a is a shadow photograph using the type 54 Polaroid film of the liquid jet under partial breakdown conditions and Fig. 9b shows a self-illuminated image of the corona discharge, using the type 57 Polaroid film, under reduced lighting and an exposure time of 15 minutes. It is observed that the highest light intensity occurs where the jet emerges from the orifice and this reduces along the jet and along the nozzle wall. By comparing the two images it is clear that the air at the orifice surrounding the liquid jet breaks down, however whether this is precipitated by electrical breakdown of the space charge laden liquid jet is unclear. The corona provides a path for charge to transfer from the liquid jet to the earthed nozzle body via the air, thus reducing the spray current and increasing the leakage current. The fraction of the useful spray current lost to earth via the surface corona discharge may be influenced both by the mean velocity, as discussed below, and the orifice diameter.

C. Overview of the Breakdown Regimes

To summarize, in the sub-critical regime, a complete electrical breakdown occurs due to electrical failure of the liquid inside the nozzle. In the super-critical regime, a corona discharge occurs outside the nozzle. The magnitude of voltage V that can be applied before a complete electrical breakdown of the liquid itself inside the atomizer, or a partial breakdown by corona discharge of the surface charge on the jet to the earthed atomizer body is defined as the ‘critical voltage’, VC. Figure 10 shows how the critical voltage VC varies as a function of flow rate Q and orifice diameter d, and also highlights the domains where ‘sub-critical’ and ‘super-critical’ atomizer operation occurs. Generally, VC increased with Q and reduction in d for the version 2 and 3 nozzle designs. The value of VC for d = 140 μm is similar with d = 116 μm for the version 3 design, implying that the electric field at the needle tip E, and hence VC, is dependent on the tip radius38 rp. The regime transition was found to approximately coincide with a mean velocity in the orifice channel of around uinj ≈ 6 m/s.

IV. Total Current versus Voltage: Observations

The variation of the total (injected) current, IT, as a function of the applied voltage, V, for different Li/d

ratios for d = 116, 140 and 254 μm at uinj ≈ 10 m/s, for a ‘version 3’ nozzle design (Fig. 2) using Diesel oil is shown in Fig. 11a-c, respectively. Results in Fig. 11d using the version 2 nozzle design and kerosene as a test liquid for d = 250 μm and uinj ≈ 10 m/s are included to quantify the effect of liquid physical properties. The termination of the curves represents the final measurement point where a partial breakdown event occurs within 1 kV of the critical voltage VC, or at the maximum applied voltage of Vmax = 24 kV if the partial breakdown did not occur, as in the case for d = 254 μm at Li/d > 3.1.

The shape of the IT-V response is similar for Fig. 11a-d and a conclusion is that the electrode gap has much more influence on the charge injection law than the orifice diameter.

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A. Total Current versus Voltage: Comparison to Quiescent Fluid Data

For quiescent liquids two regimes of charge injection processes were observed38 and later confirmed by Bonifaci et al.39 for point electrodes held at a negative polarity, and the same is observed here.

A relationship of the variation of IT with V may be developed of the form

( nT VVI 0A −= ) (1)

where A, n and V0 are constants. A is a function of the geometry (and the fluid), and V0 is a ‘threshold voltage’ that delineates the two regimes of charge injection processes (not the ‘sub-critical’ and ‘super-critical’ atomizer operating regimes) .

The threshold voltage V0 appears to occur at V ≈ 1.9, 2.9 and 4.1kV for d = 254, 116 and 140 μm respectively using the version 3 nozzle design. The slight increases in the value of V0 are due to the order in which the experimental work occurred, started with d = 254 μm followed by d = 116 and 140 μm. As the experiment proceeded the needle tip gradually melted and as a result increased the size of the tip radius, rp from approximately 5 to 20 μm. For the version 2 nozzle design using kerosene as the test liquid with rp ≈ 60 μm, V0 appears to occur at V ≈ 4.0kV for d = 250 μm.

From Fig. 11c, where d = 254 μm, there seems to be two regions of dependence where the parabolic profile of IT-V is different in each case. Where Li/d is relatively large, i.e. Li/d ≥ 3.1, there is a parabolic profile of IT-V over the range 0 ≤ V ≤ 24 where n ≈ 2. This behavior has also been observed in earlier work using kerosene for similar atomizer geometry18 as shown in Fig. 11d for Li/d ≥ 3.2. The relationship in eqn. (1) has also been observed in charge injection work of other workers for quiescent liquids40, where A in eqn. (1) is described as a function of Li with an essentially quadratic variation of IT-V (i.e. n = 2). Here the medium is in motion, and for larger Li/d a relationship is apparent. For instance, taking the Li/d = 2.4 data, denoted by the (◊) data points in Fig. 11c, the relationship between IT and V is approximately parabolic, where n ≈ 2 for V<4.8kV. However, for smaller Li/d, n > 2, which is similar to the results of Higuera41 for applied voltage V well above the threshold value V0 .

2AVIT =

A regime transition38 was found to occur at a threshold voltage V0 for emitter tip radius of rp > 0.4 μm, which together with the liquid physical properties, inter-electrode distance Li and the size of rp determined the magnitude of electric field Ep at the emitter surface38 as

[ ]pipp rLr

VE

4ln2 0= . (2)

The IT ∝ V2 relationship well above the threshold value V0 with larger tip radius of rp > 0.4 μm was

reproducible38 and was also found in an earlier work by40 and later by Atten et al.42 for needle-to-plane configurations. The constant A in a ( )0A VVIT −= relationship was found to be a function of inter electrode distance Atten et al.42 with a A2 = Li

-m variation law, where m ≈ 0.7 to 1.0 for insulating transformer oil of viscosity μ = 0.024 Ns/m2. The optimum value of voltage was found to be a function of liquid viscosity43, electrode geometry44 and polarity45, and inter electrode distance46. The breakdown processes lead to the melting of the needle material47, and avoiding the complete electrical breakdown meant that the value of the applied voltage in the present investigation was limited to the critical value VC.

The needle tip radius was rp ≈ 5 μm ± 1 μm for the work discussed here12 compared to rp < 1 μm for the specialized UO2 emitter material in the “Spray Triode”16 and rp ≈ 60 μm in the case of Shrimpton & Yule20. Figure 12 shows the electron micrograph of a used needle tip, obtained using a JEOL Model JSM 5300 scanning microscope with a 350× magnification at the end of the experimental work. The electrode surface (i.e. in the ring) seems to reveal melting processes at the needle tip, which increased rp from approximately 5 μm before the start of the experimental work to approximately 20 μm towards the end. Breakdown processes with high injected currents IT are believed to lead to the melting of the needle material48.

However, the size of rp is known only to affect the IT-V characteristic38 and the value of the threshold voltage V0 as previously. We assume it does not have an effect on the critical (i.e. optimum) atomizer operation point as the applied voltage is well above the threshold voltage V0. The critical point here is defined as the point where VC is well above V0 but below a certain value at a breakdown or discharge event. The optimum value VC is defined as 1 kV before a complete electrical breakdown occurs (i.e. in the sub-critical regime), or before the charges in the liquid jet induce corona in the surrounding gas (i.e. in the super-critical flow regime).

The derivation of a A2 = Li-m variation law was compared with the result of Atten et al.42 who used

insulating transformer oil of viscosity approximately 10× higher than diesel oil and similar tip radius size of rp ≈

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3 μm. Our rate of IT increase with V is predicted to be higher than that from Atten et al.42 as the applied bulk convection in the present study helps to strip ions from the charged layer in the vicinity of the electrode tip especially at small Li/d ratios. Small Li/d and small d conditions are more efficient because the needle tip, the charge injection site, has been translated into a region of faster moving liquid. Thus, the comparison was made for data obtained at a low applied bulk flow and at Li/d ratios bigger than 1.0. Figure 13a shows an I-V relationship for the version 3 nozzle design using d = 254 μm with 1.0 ≤ Li/d ≤ 4.9 μm at a low and steady liquid injection velocity of uinj ≈ 6 m/s, up to the maximum IT. Large relative errors at low voltages, V < 4 kV, i.e. below the threshold voltage, also make it unlikely for the data points to be included in this analysis. When subtracting the low Ohmic current from IT, the variation of IT = A2(V-V0)2 is shown by the straight lines in Fig. 13b. The ( 0A VVIT −= ) relationship was used by Atten et al.42 for analyzing the variations of IT above the threshold voltage V0.

By plotting A2 as a function of Li, Fig. 14, a continuous line with a L-0.6 variation for Li < 800 μm and a dashed line with a L-1.2 variation for Li > 800 μm were found. For comparison, Atten et al.42 reported L-1 and L-

0.7 variations for L < 20 mm and L > 20 mm respectively for 5 ≤ Li ≤ 60 mm in a quiescent liquid of higher viscosity. The differences in the correlation may be due to a laminar-turbulent transition of the EHD plumes as suggested by Atten et al.42. A refined experimental investigation using more viscous liquids with the charge injection atomizer is required in order to confirm the variation laws.

V. Effect of Flow-Rate/Injection Velocity

The variation of injection current IT with V for different volume flow rate Q is briefly investigated here to further elucidate the empirical charge injection law for V > V0, particularly for small Li/d ratios, where the atomizer is most efficient, as a function of applied bulk flow. Figure 15a-c shows IT versus V for different Q at Li/d = 1.0 using the version 3 nozzle design with diesel oil as the test liquid for d = 116, 140 and 254 μm respectively. Results from Shrimpton18 in Fig. 15d using the version 2 nozzle design and kerosene as a test liquid for d = 150 μm and for different Q at Li/d = 1.0 is also shown here for studying the effect of liquid physical properties on the charge injection process.

The general trend of the graphs shows that the rate of IT rise with V increases and is almost independent of applied flow rate Q and weakly dependent on liquid physical properties. Higher magnitudes of maximum IT are obtainable at higher Q, which is simply understood due to higher magnitudes of VC and increases the ‘flushing’ velocity.

VI. Spray Specific Charge

Figure 16 shows the spray volumetric specific charge qv=IS/Q, plotted against voltage. There is a tendency to exhibit self similarity for the two groups of flow rates. Three curves are plotted showing the sub-critical and two ‘families’ of super-critical regimes. As noted above, in the super-critical regime there is a maximum followed by a sharp reduction, and the sub-critical regime curves are truncated where catastrophic breakdown occurs inside the nozzle. The profiles in the super-critical regime, for Q≥1.25ml/s after partial breakdown has occurred exhibit similar shapes and all show that the charge remaining within the spray is now independent of both flow rate and applied voltage. There is a trend for the optimum operating points, defined as when the spray contains its maximum charge before breakdown to occur at higher applied potentials for higher flow rates. For the super-critical regime two sub-groups of data are apparent, one, with flow rates Q ≥2.27ml/s exhibit reduced spray specific charges relative to other sub-group at lower voltages. For Q =2.27ml/s, Re≈4000 (based on the orifice channel) and turbulent flow will become increasingly prevalent between the entrance of the orifice channel and the atomizer exit. Due to the existence of an additional turbulent transport mechanism for charge for the higher flow rate ranges it is thought that the radial transport of charged particles away from the centerline will be enhanced. Thus a greater charge concentration will occur near the surface of the jet as it emerges from the orifice, and this will make the surface partial breakdown condition more likely and reduce the (volumetric) spray specific charge in the jet at the breakdown condition.

The use of the version 1 atomizer geometry was confined to larger orifice diameters (d≥500μm) and the version 2 atomizer was used with kerosene for a complete data set of flow rates and diameters using the fixed geometry ratios Li/d=1.0 and L0/d=2. Changing to the version 2 atomizer while keeping d=500μm raised the spray to leakage current ratio for any flow rate and increased the spray current by approximately 15% relative to the version 1 nozzle. The trends of variations in maximum spray charge with orifice diameter and flow rate described below are typical of both version 1 and 2 designs.

In Fig. 17 the maximum spray specific charge, for each flow rate and diameter of the version 2 atomizer, is plotted against Reynolds number for kerosene. As shown, significant improvements in the maximum spray specific charge can be made by using smaller orifice diameters. These improvements can be further magnified

7 American Institute of Aeronautics and Astronautics

by using larger flow rates to a much greater extent than for the larger orifice diameters. This trend has been found to be common to the other test liquids with the added factor that more viscous, denser hydrocarbons were able to contain larger maximum specific charges. As noted above, this is assumed to be caused by the reduction in ion mobility at increased viscosity which reduces the electric component of radial charge transport. The greater efficiency of the charge injection method for more viscous insulating liquids makes the charge injection atomization technique eminently suitable for such liquids.

The reason for the reduction in the spray specific charge shown in Fig. 17 for the highest flow rate and the d=150μm diameter nozzle is hydraulic flip, the mean injection velocity being 37m/s through an orifice with no inlet chamfer. This is supported by experimental observation that the spray angle was markedly narrower for this operating condition. The inter-relationship between spray specific charge and orifice diameter may be clarified by deriving an approximate expression for the radial electric field with the liquid jet where it emerges from the nozzle orifice. The Poisson equation represents the relationship between space charge and the radial electric field in the orifice channel,

1 vqVrr r r ε

∂ ∂⎛ ⎞ = −⎜ ⎟∂ ∂⎝ ⎠ (3)

If the radial profile of the specific charge in the orifice channel is assumed to be uniform, then the following solution exists for boundary conditions Vr=0 is finite and Vr=d/2=0, representing the earthed nozzle wall just prior to the emergence of the liquid jet.

ε4rqE v

r = (4)

4rqD v

r = (5)

Equations (4) and (5) give the radial component of the electric (E) and displacement (D) fields due to a uniform radial space charge. The calculated radial electric field strength at r=d/2 is plotted against Reynolds number in Fig. 18 and as shown the data collapses quite satisfactorily. The numerator of the right hand side of eqns. (4) and (5) is roughly constant and has units C/m2, thus relates to displacement field, D, or surface charge density. It is this quantity that limits the maximum spray specific charge in the super-critical regime. From the dependence of spray specific charge on the liquid injection velocity it is deduced that the increased flow velocity delays the build up of charge near the jet surface and the partial breakdown event. The data distributions in Fig. 18 have been given linear fits and these indicate that two relationships apply and that the boundary between these occurs when Re≈4000. This agrees with arguments made above concerning Fig. 16 and it may be concluded that there is indeed a laminar/turbulent transition at Re≈4000 and that the turbulent transport mechanism detrimentally affects the maximum spray charge that the spray can contain. These linear curve-fits give the following empirical relationships between spray specific charge and Reynolds number for the version 2 design where Li/d=1,

( ) 4000Re:Re2200105.3 6 ≥+×=r

qVε (6)

( ) 4000Re:Re4200108 6 ≤+×=r

qVε

The results obtained in Fig. 18 apply to the radial component of a cylindrical electric field in an insulator of permittivity 2.2 at r=d/2 where the radial electric field is non-linear at the jet surface due to the cylindrical geometry. A commonly quoted electrical breakdown strength for air49 is 3×106V/m for atmospheric pressure and planar geometry and this compares closely with the intercept at small Re on Fig. 18. The similarity of these values are probably co-incidental however since correlations are available for the surface electric field required for corona onset of a cylindrical conductor of radius r, of which the best known is that proposed by Peek49 and for air at 760mmHg and 20°C,

)/(308.0131 5.0 cmkVr

E r ⎟⎠⎞

⎜⎝⎛ += (7)

Clearly the satisfactorily collapse of the data shown in Fig. 18 implies that eqn. (7) is not applicable in the present case since it predicts a limiting surface electric field proportional to r –1/2, where it is shown from Fig. 18 the limiting field is proportional to r. As shown by the data this hypothesis does not apply to charge carrying dielectrics.

The properties of Diesel oil used in Bankston30 are assumed to be similar to that of the fluid used in the present work (i.e. Diesel No. 2). The properties of Kerosene oil used in reference50 are assumed to be similar to that used in Shrimpton18. The ranges for viscosity, density and relative permittivity are 0.0009 ≤ μ ≤ 0.030 Ns/m2, 780 ≤ ρ ≤ 850 kg/m3 and 2.0 ≤ εr ≤ 2.2 respectively. The critical spray specific charge qv versus Re for various data obtained using charge injection atomizers for various fuels and sources as summarized in table 2.

8 American Institute of Aeronautics and Astronautics

The critical specific charge qv is enhanced for smaller orifice diameter d as shown in Fig. 19. The results also show that the magnitude of qv for the version 3 nozzle design using Diesel oil as a test liquid is higher than that of the version 2 design using kerosene for similar d in the super-critical regime of operation. This confirms that the spray, generated by a charge injection atomizer, operating in this regime holds more electric charge if the liquid viscosity is larger. Overall the dominance of orifice diameter as the primary parameter is shown.

VII. Beyond the Point-Plane Atomizer Concept

Two important problems arise with the point-plane atomizer concept described above, where the tip of the high voltage electrode (the ‘point’) and the atomizer orifice (the ‘plane’) exist with a separation of Li/d~1. The first is that as d reduces so does Li, and the precise coaxial alignment required between electrode tip and orifice centre becomes increasingly difficult to achieve. The second is that with the point plane concept, multi-orifice atomizers are difficult to construct in order to maintain a uniform hole to hole mass and charge flow. Therefore, there is an unavoidable compromise between large mass flow (requiring a large orifice diameter) and highly charged sprays (requiring a small orifice diameter). Whilst the work based around the point plane concept outlined above is useful from the viewpoint of understanding the physical basis for their operation, within the context of a forced flow extension of the fundamental studies outlined in section 3, there is limited practical application of this particular method.

There have been a number of developments to move beyond the point plane concept and improve the applicability of the charge injection method to practical systems. The developments are now discussed.

For a reasonably sized orifice diameter atomizer employing the point plane concept a limitation, is that a ‘two zone’ spray tends to be generated. Typically a spray plume consists of a poorly atomized and charged ‘core’ surrounded by a highly charged ‘sheath’ of smaller more highly charged drops. One possible solution to this problem is to employ a secondary atomization method and the fusion of a pressure swirl atomizer and a charge injection atomizer has been suggested using a number of strategies.

Kelly16, whilst at Exxon, reported the ‘Spray Triode’ modified to operate as a pressure swirl atomizer but gave very few details of the design and performance. It was claimed the addition of an imposed potential of 7 kV changes the distribution, disrupting all drops where D>150μm. The reported drop size distribution for charged spray atomization shows an absence of drops larger than approximately 120μm but the distribution appears truncated. The drop size measurement or sample size is not quoted, nor are any details of the device.

Laryea & No51-52 took a standard pressure swirl nozzle and inserted a high voltage electrode along the axis. They reported a significant improvement for their application of crop spraying, but the coincidence of the high voltage electrode and the air core of the pressure swirl atomizer may have hindered the performance.

Anderson et al.53-55 modified a gasoline direct injector (Mitsubishi p/n MR560552) to inject electrostatic charge into the fuel by externally mounting an exposed high voltage electrode, featuring a sharp conical tip at the orifice. This electrode made from copper, had a sharp cone at the tip of the electrode protruding 0.5 mm at an angle of 135 degrees with the base. The locally intense electric field that is generated from the sharp conical feature partially polarizes the fuel molecules and injects ions into it, thus transferring a net charge to the liquid. The degree of spray charging appears not to be significant enough to produce meaningful changes in the primary atomization nor spray dispersion. In addition, in contrast to charge injection atomizers, the external exposed high voltage electrode is vulnerable to corona discharge in an ionized combustion environment.

Here the aim is to obtain high volume flow rates at low pressure drop with a finely atomized and highly charged spray. In the context of electrostatic methods based around the charge injection concept, the only possible solution is to multiplex the number of orifices present. The engineering challenge then moves to ensuring each liquid jet emerging from each orifice carries a similar amount of current as its companions. Two technologies have emerged in the commercial sector, one of which remains protected by patent, the other has been released to to public domain. Kelly56-59 suggests a number of possible design concepts and two SAE papers explore this60-61. Similarly Allen62-63 suggests some concepts and these have also been evaluated64-65, which are now in the public domain

Here the typical application is fuel injection for internal combustion spark or compression ignition engines. There are a number of reasons for the interest. For compression ignition engines, electric charge in the drops will aid dispersion and mixing, and can possibly improve the combustion emissions, especially NOX. For direct injection spark ignition engines there is the possibility of controlling the spray more effectively moving from an early to a late injection strategy, since electric spray charge is in effect a free parameter. There is also interest in using EHD injectors to enable the downsizing of engines for enhanced fuel economy, for instance within hybrid systems.

In this scenario atomizer design and operation is particularly challenging since the flow and indeed the atomizer geometry are time dependent, and the high voltage electrode must be activated correctly. In addition various time delays must be accounted for (start/end of fluid motion, voltage rise time, time for charge to be injected, needle lift response).

9 American Institute of Aeronautics and Astronautics

Kelly, in collaboration with Robert Bosch GmbH, patented a EHD single hole CI injector design66 and Bosch published some performance data67. Experiments were carried out and some tests resulted in very high specific charges (up to 5C/m3) and consequently very small drop sizes (<50μm). Experiments were also conducted for pulsed voltages of several milliseconds. Time delays were apparent, allowing for the voltage rise time (~100μs) and also the time taken for the spray current to rise from its initial to its final value (charge rise time), allowing for the voltage rise time. Typical delays of several 100μs correspond to charge mobility velocities of 1 m/s, correct for the hydrocarbon fuel tested. Increasing the electrode gap increases the time taken for the current to start to rise and increasing the flow rate reduces the charge rise time. The onset of corona discharge typically shifts to higher voltages for shorter duration pulses. Charge density measurements were also found to be higher than for stationary conditions. They note that at extreme conditions of high voltage, short pulse operation leads to as yet undefined time delay effects which limit operation. Allen68 has patented a multi-hole pulsed EHD atomizer though no data is available, and this is now in the public domain.

There are a number of other developments in the subject, and all are patented by Kelly. Of most interest is an ‘electron gun’ system, known as the ‘Spraytron’, a photo may be found in Kelly34, Figure 11 of that paper, and the relevant patents are Kelly69-72. In this concept the pointed metal high voltage electrode is replaced with an electron gun and an ‘electron transparent window’. Other patents relate to electronic control systems73-75.

VIII. Spray Characterization and Combustion

For charge injection atomization, of electrically insulating liquids, it is typical that a polydisperse drop diameter distribution is present in the spray. It is also typical that the mean trajectories of the drops are a function of the mean drop diameter, and that the smaller drops tend to be found on the spray periphery, and the larger drops nearer the spray axis. These observations are now discussed in more detail. A. Spray Visualization and Prediction of Expansion Rate

Spray visualization is described for the three cases described in Table 2. It is noted that an uncharged 'spray' from the atomizer, at the conditions described in Table 2, consists of stream of large drops moving along the spray centerline. Figure 20 shows the case for a flow rate of 0.5mL/s and a specific charge of qv=-1.20C/m3 and it is seen that the plume expands significantly. This figure highlights the dual-zone nature of these particular charged sprays. Recirculation is also observed near the injector, and this is due to small (D≈5μm), highly charged drops, produced from jet break-up, being attracted back towards the earthed atomizer body. For higher flow rates and specific charge, as shown in Fig. 21, the amount of scattered light increases significantly which implies higher drop concentrations. The spray, under the influence of space charge forces, expands to fill the volume illuminated. It is suggested that the axial drop velocity components are generated primarily by the jet flow, while the radial deflections shown in the figures are caused by electrical forces acting on the drop trajectories. This proposition is supported by CFD modeling of this type of spray76 which revealed that the radial electric field, due the approximate cylindrical symmetry, is at least an order of magnitude greater than the axial electric field generated by the charged drops near the centerline. The drop trajectories are very well ordered, typically emerging normally from the spray core and diverging at larger radial displacements.

Based on the method developed by True77, Shrimpton & Yule78 developed a simple force balance model that was sufficient to explain the interplay between axial momentum and radial dispersion. This is achieved using the following simplifying assumptions: 1) Monosized charged spray, where drop and orifice diameter are equivalent. 2) Axial velocity changes are due to aerodynamic drag forces only. 3) Radial velocity changes are due to space charge forces and drag forces. 4) The gas phase is quiescent. 5) There is no drop evaporation or break up. The axisymmetric charged spray plume was modeled as a sequence of cylinders, each containing the same mass of drops, each of charge q, which is uniformly spread throughout the volume of the cylinder. This assumption allows an analytic solution of the radial electric field. Bearing in mind the above assumptions, the equations of motion for the axial (u) and radial (v) velocity components may be written,

u|u|8D C =

dtdu

m xxg

2

Dx ρ

π (8)

u|u|8D

C + E q - = dtud m rrgDr

r ρπ 2

(9)

10 American Institute of Aeronautics and Astronautics

where m, q and D are the drop mass, charge and diameter, and ρg is the gas density. The radial component of the space charge field generated by the charge on the drops within the spray plume, Er was obtained by applying Gauss's Law to a cylindrical volume containing an average volumetric charge qv, such that

r 2q

= E0

lr

επ (10)

where ql is simply the charge per unit length, r is the radius of the volume. The drag coefficient used is a standard empirical form79,

⎟⎠⎞

⎜⎝⎛ += 3/2Re

611

Re24

dd

DC (11)

where Red is the drop Reynolds Number. A model jet break-up length, lj may be defined from defining characteristic velocity and break-up times78 and initial conditions are typically defined from global nozzle parameters:

D=d ; ux,o=uinj ; ur,0=0 ; q = qv 6D 3π

(12)

The trajectory of the modeled spray plume boundary, considering the simple form of the model, is representative of the experimental data. Evaluation of the model78 showed that the simplistic model is successful in capturing the evolution of spray shape as a function of the initial conditions. This shows that the axial motion is governed by drop momentum and aerodynamic drag and the radial motion by an approximate balance between electrostatic and aerodynamic forces.

It has been shown that for highly charged sprays (qv≥1C/m3) drop deflection away from the central core is significant and for efficient combustion of these charged plumes some form of spray confinement may be necessary. Disruption of the spray core/sheath boundary is shown in Fig. 22 for the case of a flow rate of Q=0.5mL/s and a spray specific charge of qV=1.20C/m3 for an orifice diameter of d=250μm. The disruption is caused by the placement of an earthed bar of diameter 12mm placed at x=0.08m, y=0.08mm, in addition to earthed surfaces along x=y=0.15m. The attractive force felt by a charged particle to a flat earthed plate (the 'image' force) is only a quarter of the force of attraction of two oppositely but equally charged particles acting over the same distance. This ensures that space charge forces, generated by the charge on the drops, will be dominant except in the region near the earthed surface.

The preliminary test shown in Fig. 22 demonstrated the effect of altering the electric field at discrete regions rather than planes and this principle is now investigated for a system of ring electrodes. The ring electrodes were 100mm in diameter and fabricated from 1/4" (6.25mm) steel bar. They were placed symmetrically along the axial centerline of the spray and results are shown for the same nozzle diameter (d=250μm), flow rate of Q=1.6mL/s and spray specific charge qv=1.80C/m3. Figure 23 shows the spray for both electrodes earthed. In the region bounded by the nozzle, the upper electrode and the plume boundary, the small drops follow the field lines emanating normally from the spray plume boundary to either the earthed nozzle electrodes or the earthed upper ring electrode. The illuminated region is distinctly bounded by a distorted quadrant. Similar behavior was observed underneath the upper electrode and in the vicinity of the lower electrode. However for these positions the further away the drop is from the earthed electrode the less is the magnitude of the image force, and drop momentum thus has a greater influence on the drop trajectory. This is not observed for the region above the upper electrode since the drops are very small and quickly obtain a terminal velocity from the interaction of drag and electrical forces. The zero normal gradient that exists along the horizontal line halfway between the two electrodes and away from the spray core is observed. The essentially symmetrical behavior of the drops near the electrodes above and below the symmetry line shows that the image forces are dominant in the near electrode region. A fraction of the drops in the plume recirculate underneath the lower electrode and ultimately reverse their trajectories. Drops that have recirculated and possess radial displacements greater than the ring electrode radius (50mm) are either attracted or repelled from the electrode and this is understood by considering image and space charge forces. Drops near the rings follow the field lines and impinge onto the electrodes. For drops further away from the rings, where there are additional charged drops between the drops considered and the ring electrodes, the space charge repulsion between the drops dominates over the image charge force and the drops are repelled away from the electrodes. For drop trajectories in the diffuse outer region the tracks show a 'wiggle'. It is unclear whether this is due to changes in drop shape, and hence the centre of mass, or, due to velocity fluctuations caused by the non-uniformity of the space charge field in the diffuse region.

The effect of applying positive and negative potentials to the ring electrodes on the spray dynamics is now presented. Figure 24 shows the case where the potentials of the upper and lower electrodes are positive and negative and have a magnitude of 7kV. The flow conditions for Fig. 24 are a flow rate of Q=30mL/min and a spray specific charge of qv=1.20C/m3. The large recirculation caused first by the additional radial deflection from the interaction of the upper positive electrode, and secondly by the repulsive force from the interaction of

11 American Institute of Aeronautics and Astronautics

the negatively charged lower electrode, is similar for both potentials. Higher drop numbers in the recirculation zone were observed for the higher potential case, and similar behavior was observed for higher flow rates and specific charges. Figure 25 shows the case for the same flow conditions and the same magnitudes of the potentials as Fig. 24 respectively, but with reversed polarity of the rings. Above the upper electrode the small drops populating this region are deflected upwards, away from the like charged ring and towards the earthed nozzle. Once drops are at axial displacements greater than the upper electrode they are attracted towards the lower electrode. However it is noted that in the region between the lower electrode and the injector, the spray is effectively confined.

This has importance for the possible application of electrostatic atomization of hydrocarbon oils to spray combustion systems, where the divergent plumes could lead to incomplete combustion of the spray plume and the possibility of tailoring the spray shape may offer advantages and solutions. B. Quantitative Spray Characteristics

Predominantly this summary is taken from references18-19,78,80-81. Phase Doppler Anemometry (PDA) results are now presented for cases 1-5 in table 3. Figure 26 shows the drop diameter frequency distributions at a position 0.12m downstream of the atomizers on the spray axis, where the photographic results showed that the majority of the ligament structures had disappeared for all cases. In terms of a representative position for the spray, the axis represents the “worst case” in terms of atomization quality because a high proportion of the more highly charged smaller drops have migrated away from the central core. The drop diameter frequency distribution, for the 500μm orifice diameter nozzle (cases 1-3 of table 3), is very different from the other cases. Figure 26a,b-c show two distinct drop populations. The population of very small drops, D<5μm, was found throughout the spray plume. It is thought to be generated during the primary atomization of the jet by a Rayleigh–type break-up of charged ligaments. The small drop end of the spectrum also includes drops in the size range 20-40μm, and these drops are thought to be the product of secondary atomization of the larger droplets. At the other extreme of the size range there is a population of large drops with the peak of the distribution at approximately 440μm for the lowest flow rate, case 1 of table 3. The position of the peak decreases slowly as the flow rate rises. These large droplets are the product of the primary atomization of the relatively poorly charged spray core. There is a low peak slightly to the right of this second distribution. Although small on a number basis, the volume of the large drops in this subsidiary peak equates to approximately twice the volume of smallest drops. Instability in the jet core is thus producing drops of diameter slightly less than that of the orifice. This contrasts with the drop diameter for uncharged jet break-up predicted by Rayleigh, D=1.89d. In general the low flow rate, low specific charge atomization (case 1) generates comparable numbers of very large and very small drops but few intermediate sized drops. As the flow rate is increased, more small drops of diameter range 20-40μm are generated. Given that the specific charge for cases 2 and 3 is identical, the additional secondary atomization for case 3 (giving more drops between 20μm and 40μm) must be at least partially aerodynamically driven, and caused by the higher nozzle velocity for this case. The nature of the resultant drop diameter distribution in case 3 is unlike the products of uncharged bag/strip secondary atomization modes, where the resultant drops have a much greater range of diameters82. Therefore there exists a secondary atomization mechanism which is particular to charged drops, but is also a function of aerodynamic effects.

There are similarities between the sprays of cases 1, 2 and 3 of table 3, generated via a charge injection mechanism, and the ‘electrosprays’ of doped heptane7. Both data-sets possess a generic dual-zone spray character, with a bi-modal drop diameter frequency distribution with the ratio of diameters between the large and small drop populations being approximately 10:1 in each case. Gomez & Tang83 presented photographs of the secondary atomization of the large primary drops, which showed this value for the ratio of drop sizes before and after the process. Thus it is likely that this mechanism is responsible for the 20≥D≥40μm size class shown in Fig. 26b and Fig. 26c. It seems that despite two very different methods of charging the liquids, and the absence of doping to decrease the liquid resistivity in the present case, the sprays are remarkably similar. This is despite a three order of magnitude difference in resistivity between doped heptane7 and the kerosene of the present work. A significant difference however is that of the sizes of the largest drops. In the case of the present hydrocarbon sprays, the large drops for cases 1, 2 and 3 of table 3 are of the order of the orifice diameter (500μm) while for the charged heptane sprays, the largest diameters were approximately 40μm7. This suggests that a change in liquid electrical resistivity is primarily relevant to the liquid charging process, and different ranges of resistivity are required for different electrostatic atomization techniques, as discussed in the introduction. However once the liquid contains charge, added by either the cone-jet or the charge injection method, apart from the scale consideration, the spray structures are similar.

There is an absolute limit regarding the performance of liquid atomization via a Taylor-Cone. For a given atomizer geometry, an increase in flow rate will generally produce an increase in both the mean diameter and the breadth of the drop diameter frequency distribution if a Taylor cone is used4. In the present work it is found that this is not the case for sprays generated using the charge injection technique, since the maximum spray

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specific charge attainable generally increases with flow rate. This was found to be most apparent for charge injection atomizers where d≤250μm, for example for cases 4 and 5 for which drop diameter distributions at x=0.12m, r=0m are given in Fig. 26d-e respectively. These show that for the higher flow rate, case 5 of table 3, the distribution is skewed more towards the smaller drops and the most frequent drop diameter is D≈140μm compared to D≈250μm for case 4 of table 3 Fig. 26d. The maximum specific charge achieved increases with flow rate, thus this shows that promotion of liquid atomization occurs via the combination of increased aerodynamic and increased electrostatic forces. This is the opposite behavior to standard contact and induction charging techniques and is the reason that charge injection techniques are so well suited to high flow rate operation. Comparing the drop diameter frequency distribution results for the lower (cases 1, 2 and 3 of table 3) and Fig. 26a-c and higher specific charge sprays (cases 4 and 5 of table 3) and Fig. 26d-e reveals that in the latter cases the frequency distributions are continuous. This behavior cannot be due to a laminar/turbulent transition occurring within the nozzle leading to enhanced atomization for some cases. This is because the Reynolds number, based on conditions at the atomizer orifice for case 2 is double that of case 4. Superficially there seems to be no obvious linking trend between the PDA data of cases 4 and 5, shown in Fig. 26d-e, and either, (1) the low flow rate PDA data of cases 1, 2 and 3, shown in Fig. 26a-c) or, (2) the similarity with electrosprays noted above. However electrosprays with continuous drop diameter frequency distributions have been reported in the literature84-85 where ethanol was sprayed at a high flow rate, 0.06ml/s, from a capillary held at a very high applied potential, typically 20kV. In this case the atomization mode has been characterized as ‘rim-emission’86. Thus cases 4 and 5 of table 3, the sprays of Grace were obtained by operating at relatively high liquid throughput and at a high spray specific charge for the respective atomization methods. In contrast to the work of Gomez & Tang7,83, Grace & Dunn85 observed drop generation from ligaments emanating from the electrified meniscus rather than a single jet as in the case of a Taylor Cone, and this produced a polydisperse spray, with a maximum diameter of approximately 40μm. A similarity of the spray structure and drop diameter frequency distribution at high charge densities (cases 4 and 5 of table 3) with the results of Grace is apparent. In turn, a similarity exists for the low charge density results and the work of Gomez & Tang7 and others. Thus the classification of sprays as either a well-ordered dual zone character or a more chaotic polydisperse form, occurs in both semi-conducting liquids, atomized by applying an electrical potential to metal capillary, and highly insulating liquids, atomized using a charge injection technique.

A further insight into the spray character may be obtained by examining the variation of mean drop diameter along the axis of these cylindrically symmetric sprays. Figure 27 shows contours of numerical mean diameter D10 obtained by processing the PDA data for cases 2, 4 and 5 of table 3. The stratification of drop diameter with radial displacement is clear and this is more marked and different than for other atomization methods. For example for full cone and hollow cone pressure jet sprays, and two-fluid atomized sprays, there is a tendency for smaller droplets to occur at the spray centre. Since all drops are injected with a negligible initial radial velocity component the radial stratification of drop diameter must mainly be due to electrostatic forces. The sprays are steady-state, so that the electric field is approximately constant, the only variations being due to local space charge fluctuations. The main parameter that may affect the trajectory of a charged drop is thus its specific charge. Since a radial stratification of drop diameter occurs, the charge to mass ratio is unlikely to be constant and the trend is explicable by smaller drops having larger charge to mass ratios. The variation in mean diameter along the spray axis is due not only to smaller, more highly charged drops moving away from the axis, but also to larger drops becoming “visible” to the PDA when they attain the required sphericity. For case 2, shown in Fig. 27a, the boundary of the inner core of large drops is clearly visible, showing the dual-zone spray characteristic, while for the more highly charged sprays this boundary becomes blurred and gradients are more gradual. Figure 26-27, and the above discussion, shows that the atomization mechanism for producing drop diameter distributions for these charged sprays is a combination of aerodynamic forces and the disruptive electrical forces caused by the presence of free surface charge. The specific charge attainable increases with injection velocity, for a all orifice diameters. Therefore the combination of larger electrical forces that larger higher aerodynamic shear at larger injection velocities improves atomization performance significantly. As described in a section 7.1, these charged, well atomized sprays permit a degree of manipulation and control by the application of additional applied electric fields. They are also ignitable such that stable combustion can be achieved. It is also worth noting that since the charge to mass ratios of drops within the polydisperse spray are size dependant, the possibility of essentially sorting drops by size classes is available.

From the spray visualization and drop diameter results it has been shown that although a range of drop sizes is present, the trajectories vary gradually with respect to the spatial position in the spray. Figure 28a-b show, respectively, the mean axial and radial velocity profiles for the cases of Q=0.5cc/s, qv=1.20C/m3, and Q=1.67mL/s, qv=1.80C/m3. The general pattern of spray behavior is similar to that first observed by Shrimpton et al.18 for the 500μm orifice diameter atomizer. Thus high axial velocity components and large radial accelerations occur in the spray core, even though the drop distribution is continuously varying, and radically different from the lower charged sprays.

13 American Institute of Aeronautics and Astronautics

An overview of the mean axial velocity profiles Fig. 29b shows that the axial velocity, in general, decreases with increasing axial displacement from the atomizer. This suggests that the initial momentum, from the liquid injection at the atomizer, is transferred to the gas by drag forces and this process predominates over any axial acceleration due to the electric field. Indeed axial velocity versus drop diameter correlations at particular points along the spray centerline18 have shown that the drop axial velocity is proportional to the square of drop diameter, suggesting that drag forces dominate. Of interest to charged spray dynamics is that the mean axial velocity, u=20m/s, at x=0.03m for Q=1.67cc/s is lower than the mean axial velocity at the next axial ordinate (x=0.6m) where u=24.6m/s, where, from the drag relation noted above, one would expect it to be lower. This would seem to be due to the larger drops of the diameter distribution not attaining sufficient sphericity for measurement by PDA, and biasing the mean drop velocities at small axial displacements to smaller, and hence slower drops. This behavior is not repeated for the lower flow rate case and it was not observed for sprays produced by the atomizer with an orifice diameter of 500μm. Analysis of the results from a two phase charged spray CFD code76 shows that the axial electric field caused by the charged spray changes direction between the atomizer and at an axial position 10mm below the atomizer orifice where the spray core has started to spread out. Examination of the directions of the axial electric field and sign of the drop charge showed that charged drops must be axially decelerated near the atomizer due to force between individual drop charges and the spray electric field. This experimental evidence confirms that this phenomenon does have an effect, providing sufficiently high spray specific charges are used and supports further development of the model.

Examination of the mean radial velocity profiles Fig. 28b shows similar behavior for both flow rates. In general the acceleration of the drops away from the spray centerline is higher for the spray with the larger spray specific charge. Due to the axisymmetric geometry both the mean radial electric field and also the drop radial velocity component at the spray centerline must be zero. However near the centerline, relative to the spray width, the radial electric field is a maximum and decreases rapidly as radial displacement increases. The radial velocity profiles shown in Fig. 28b closely resemble in shape the radial electric field profiles generated by a charged spray CFD code76 which suggests that the radial drop motion is wholly driven by the electric field and the drop charge.

IX. Spray Combustion

Weinburg conducted research on many electrical aspects of combustion6,11,87, and recently this work has been used in the development of electrospray combustion systems for small portable power units88-90. However these doped electrospray combustion methods rely on the liquid hydrocarbon fuel being ‘doped’ with an additive to increase the electrical conductivity sufficiently to create sprays with conventional electrostatic techniques.

Charge injection techniques permit a range of hydrocarbon fuels to be burnt without doping. Combustion tests were carried out for conditions specified as cases 4 and 5 of table 3 and the flame shape was similar to the results shown in Fig. 29. This is in spite of the spray specific charge being approximately double for the smaller orifice diameter. This supports the previous suggestion that the smaller drops evaporate and burn by the time the flame front is reached. For the case 6 combustion test, a ceramic disk was fitted onto the face of the atomizer, since cold spray recirculation and plume expansion was excessive near the atomizer and this initially raised safety concerns. On ignition however it was found the excessive spray expansion is negated and stable flame was obtained, that did not require the pilot flame for stabilization as shown in Fig. 30a. The base of the flame front was 9 to 10mm above the surface of the atomizer, as shown by an enlarged image of the flame seat Fig. 30b. Visual examination of the base of the flame front showed that the inside and outside surfaces of the flame front were surrounded by a blue halo and the region where jet break up was occurring was not burning. It is thought that the combustion and atomization zones are segregated by a fuel vapor-rich evaporation zone. These observations are similar to those of Gomez and Chen88 made for much smaller liquid throughputs using a contact charging method with doped heptane. There is an added advantage of the charge injection design used here, compared to the contact charging method, in that the high voltage electrode is enclosed within the earthed nozzle body, which improves the robustness of the burning spray flow by completely decoupling the charge injection and combustion processes.

Diesel oil was also successfully burnt using a 250μm orifice for operating conditions specified in case 7 of table 3, and this is shown in Fig. 31. This result highlights the applicability of the method for combustion of heavier hydrocarbons fuels and the improvements with respect earlier results30 where a high temperature and pressure version of the 'Spray Triode' was used which had smaller orifice diameter of 175μm and reported only intermittent combustion of diesel oil.

A study was made of the effect of pre-heating the liquid fuel on the magnitude of the spray charge and this was negligible over the small range tested, of 295 311 KT≤ ≤ . This suggests that moderate fuel preheating would be possible for the electrostatic atomization and combustion of heavy fuel oils using a charge injection method.

14 American Institute of Aeronautics and Astronautics

The effect of an additional electric field on the flame was also investigated and an approximately uniform field was applied normal to the nominal spray/flame symmetry axis using vertical metal plates. The effect of the presence of the plates when both were earthed was initially investigated. It was found that the presence of the plates caused the spray flame envelope to expand relative to a similar free spray. This is sensible considering the increase in potential gradient caused by the zero potential planes caused by the plate surfaces. The results presented below may be understood more easily with reference to similar effects previously documented87 on uncharged homogenous flames in the presence of electric fields. To summarize, for a net uncharged flame, positive ions and free electrons are created in equal amounts. However the electrons have much larger mobilities than the positive chemi-ions and are quickly accelerated and removed by the applied electric field. This leaves the flame with a much higher concentrations of positive ions and net positive charge, which ultimately results in positively charged soot particles. These positive ions will induce body forces and motion, the ‘ionic wind’, and therefore homogenous flames tend to be attracted to sources of negative potential. Stronger electric fields will increase the rate of extraction of negative charge until it equals the rate of production, at which point the current flux and body force magnitude will no longer increase.

The electrode on the left of the spray was used with different positive and negative potentials as shown in Fig. 32 and Fig. 33 respectively. The electrode on the right of each flame was earthed, and for these results the spray defined by case 4 in table 3 was used. For the positive applied electric field, Fig. 33, spray and flame deflection is small but flame luminosity is reduced at higher applied fields. The reduction in luminosity is thought to be due to the extraction of free electrons but that the deflection was so small was unexpected. Without combustion the spray plume would be deflected to the left progressively as the applied field is increased. For this case, the presence of the ionized environment is decoupling the link between the applied field and the droplets. The application of a negative electric field Fig. 33, shows that charge remains on the drops as it can be seen that the drops are deflected by the applied field. Faint drop tracks and streaklines illuminated by combustion of the drops can be seen on the right of the flame. The flame is pulled to the left towards the negative potential sink by, it is thought, the ionic wind induced from the positively charged chemi-ions, the carbon particles. This premise was confirmed by the observation that during this experiment soot deposits were found on this electrode.

In general electrostatic effects work best on small particles with high charge to mass ratios. Ionized molecular products probably represent the upper limit in terms for particle charge to mass ratio which implies that the spray shaping, mixing, flame stabilization and optimization should be easily realizable. To have the spray essentially pre-charged as a subsidiary effect of the atomization system encourages the use of electrical flow and combustion modulation. This also raises the possibility that the residual charge in the combustion products could then be used to increase the performance of emissions control systems using robust lower power designs with no moving parts of small unit size. The use of electromagnetic forces in plasma control is very advanced and it is surprising that electrical control of combustion systems has not seen the same success. In particular the use of AC electric fields to synchronize spray placement with combustion oscillations is suggested as a way of reducing soot and NOx simultaneously.

. X. Future Outlook

Electrostatic atomization methods and electrical sprays have been widely used in industry for decades, most

notably for painting applications. In science and several high technology industry sectors, electrical sprays are well established and highly regarded, most notably the Nobel Prize for chemistry in 2002, given to John Fenn 2002 for his profound impact on the mass spectrometry industry. All of these examples however relate to sprays of electrically semi-conducting liquids, produced typically by flowing the liquid through a metal capillary held at a potential of a few kilovolts. The benefits of electrically charged sprays are well known: excellent dispersion and inter-phase mixing, absence of agglomeration, controllable spray targeting, good coating uniformity, low power consumption, good primary atomization performance.

Electrostatic atomization for electrically insulating liquids, the charge injection technique as this manuscript highlights, has been developed over the last 2 or 3 decades, so why does it not have the popularity and acceptance within the same sort of science and industry sectors that is evident for semi-conducting liquids? This section gives the interpretation of the author. The following points are made:

It cannot be denied that a charge injection atomizer is a more complicated device to manufacture and operate than a standard atomizer to do the same job. Much of the research applications of charge injection atomizers have been directed towards steady state combustion systems. These are systems which generally have large amounts of spare energy to for instance pressurize a liquid to produce a finely atomized spray, without the complexity of dealing with a high voltage electrode embedded deeply within the atomizer. Comparing electrostatic atomizers for semi-conducting and insulating liquids, the latter class of atomizer is more complex. Therefore there has to be a compelling reason for the user to accept this level of complexity.

15 American Institute of Aeronautics and Astronautics

Charge injection atomizers rely on the coupled effects of electrical and hydrodynamic forces acting together to operate effectively, and whilst this manuscript brings together a great deal of the basic physical timescales, non-dimensional numbers and empirical data on atomizer operation, these have not been fully understood. Whilst a good empirical understanding of charge injection atomizers is now available the science that underpins and explains the observations is not yet complete. In short, these are complex systems that are not completely described.

A distinct problem has been that the development of much of the early technology was undertaken in the commercial sector, and is effectively lost to the wider community for further development. Some of the theories that have been proposed to explain the spray drop-charge character have rather shaky physical foundations and have obscured the basic electrohydrodynamic coupling that is key to understanding the internal flow patterns within the charge injection atomizer and primary atomization characteristics outside it. This article is the authors attempt to move away from this purely commercial perspective. The technology, originally patented by Scion Sprays, now in the public domain is another by Jeff Allen.

Whilst laboratory work has shown the atomization process to be stable and robust on the timescale of hours, to be useful in the commercial sector, the charge injection systems in use need to be able to operate for thousands of hours between service outages, and this level of robustness is completely unproven as of present.

Despite the above we are not especially gloomy regarding the prospects of charge injection method becoming more popular, and again, a number of points, our personal opinions again, are made:

The charge injection method is the only method able to electrostatically atomize and disperse electrically insulating liquids. Therefore if an application requires excellent spray coating uniformity using an oil based solvent, this is the only option. The charge injection method is the only way to confer all the established advantages of electrically charged sprays onto electrically insulating liquids. Our view is that somewhere, sometime soon, someone will need this technology.

Several drivers are already present, and here we give a couple of examples: The food industry is one such example, for instance a bread production line. Here the tins need coating with oil before the dough is added to go on to produce bread. The spray needs to coat the tin uniformly with as little overspray and excess deposition as possible. If an electrical spray of a vegetable oil could be produced, this would provide the precise control required. Another example is small (<50cc) internal combustion engines. These currently use carburetors, and have astonishingly poor emissions characteristics. By around 2011 in the EU and the US, these engines will need to meet strict emissions that approach the level required by larger engines. Large engines have met the emissions limits by employing high pressure injection systems, providing fine atomization, good fuel-air mixing and cleaner combustion. However one cannot deploy these injection systems to small engines, since they consume too much power. Therefore one needs a low power fuel injection system to provide fine atomization and excellent fuel atomization and mixture preparation. Pulsed EHD systems are one of the very few that can provide this.

To conclude, charge injection atomizers are complex EHD systems, which are not fully understood, though this understanding continues to grow. Because of their complexity they will never be adopted where a simpler non-EHD system can perform the same task. Electrical sprays do however have distinct advantages with a high degree of control possible, and charge injection systems are the only way to produce sprays of electrically insulating liquids. Performance of spray systems, in terms of economic or environmental factors (reduction of overspray) quality control (coating uniformity) or efficiency (drop size per watt consumed by the atomizer) will produce niche technology applications where the performance advantage of charge injection systems outweighs the complexity overhead. The adoption will not be driven by choice, but it will arise from the pressures of environmental legislation and economic forces.

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16 American Institute of Aeronautics and Astronautics

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Property Unit White spirit Kerosene Diesel No.1 Diesel No.2 Mineral oil Marcol-87 Reference - [18] Measured [91] [30]

ρ kg/m3 780 800 840 815 850* 850 μ Ns/m2 0.0009 0.0011 0.0024 0.0026 0.0250 0.0300 εr - 2.2 2.2 2.2 2.2 2.0* 2.0

Table 1: Liquid physical properties

Description

Orifice diameter (d, μm)

250

250

150

Flow rate ( QL, mL/s)

0.5

1.67

0.5

Mean injection velocity (m/s)

10

34

28

Specific charge (ρV, C/m3)

-1.20

-1.80

-3.00

Reynolds Number 1500 5100 2520 Weber Number 1.2 13.9 5.6

Table 2: Atomizer operating conditions

Property Unit Case 1 Case 2 Case 3 Case 4 Case 5 Case 6 Case 7 Liquid - kerosene kerosene kerosene kerosene kerosene kerosene Diesel oil Orifice diameter

μm

500

500

500

250

250

150

250

Flow rate ml/s 1.3 2.0 3.0 0.5 1.7 0.5 1.0 Mean velocity *

m/s 6.7 10 15 10 34 28 20

Specific charge **

C/m3 0.4 0.5 0.5 1.2 1.8 3.0 1.4

Reynolds number*

- 2700 4900 5720 1900 6800 3400 1800

Table 3: Atomizer operating conditions

* Based on flow rate and orifice diameter, to 2sf. ** These are the maximum specific charge values possible for the particular orifice diameter

20 American Institute of Aeronautics and Astronautics

Figure 1: Photograph of the atomizer

Figure 2: Variation of the generic charge injection design for (a) Version 1, (b) Version 2 and (c) Version 3.

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Figure 3: Atomizer schematic.

Figure 4: A typical experimental layout and electrical connections.

Figure 5: Resistivities (σ) of the white spirit (□), kerosene (Δ) Diesel oil (◊) versus DC applied potential20.

22 American Institute of Aeronautics and Astronautics

Figure 6: Total current versus applied potential for kerosene (◊=0.42ml/s, ♦=0.63ml/s, ×=0.83ml/s), Diesel oil (∆=0.42ml/s,▲=0.63ml/s, +=0.83ml/s) and white spirit (□=0.42ml/s, ■=0.63ml/s +=0.83ml/s) in the sub-

critical regime : version 1 nozzle, Li/d=6.6, L0/d=2, d=500μm20.

Figure 7: Spray current versus applied potential for kerosene (◊=0.42ml/s, ♦=0.63ml/s, ×=0.83ml/s), Diesel oil (∆=0.42ml/s,▲=0.63ml/s, +=0.83ml/s) and white spirit (=□0.42ml/s, ■=0.63ml/s +=0.83ml/s) in the sub-

critical regime : version 1 nozzle, Li/d=6.6, L0/d=2, d=500μm20.

23 American Institute of Aeronautics and Astronautics

Figure 8: (a) Spray current in the super-critical flow rate regime using a version 1 nozzle, Li/d=6.6, L0/d=2,

d=500μm, Q=2.00ml/s for kerosene, white spirit and Diesel oil, (b) Leakage current in the super-critical flow rate regime using a version 1 nozzle, Li/d=6.6, L0/d=2, d=500μm, Q=2.00ml/s for kerosene, white spirit and

Diesel oil20.

Figure 9: Spark shadowgraph of the partial breakdown event using kerosene: version 1 nozzle, Li/d=6.6, L0/d=2,

d=500μm, Q=2.00ml/s.

Figure 10: The critical applied voltage, VC before a partial (super critical regime) or complete (sub critical regime) breakdown occurred versus applied flow rate Q at Li/d = 1.0 using kerosene and the version 2 nozzle design with rp ≈ 60 μm for d = 300 (▲), 250 (□) and 150 (♦) μm and using Diesel and the version 3 nozzle

design with rp ≈ 5 – 20 μm for d = 254 (■), 140 (0) and 116 (+) μm22.

24 American Institute of Aeronautics and Astronautics

Figure 11: Total current, IT versus the applied voltage, V at a constant injection velocity of uinj ≈ 10 m/s using

Diesel and the version 3 nozzle design with rp ≈ 5 – 20 μm for (a) d = 116 μm, (b) d = 140 μm, (c) d = 254 μm, (d) d = 250 μm22.

Figure 12: Electron micrograph of needle tip with a 350× magnification, and the measured tip radius, rp ≈ 20 ±

1 μm22.

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Figure 13: (a) IT and (b) IT1/2 versus applied voltage V at a constant injection velocity of uinj ≈ 6 m/s using Diesel

and the version 3 nozzle design with rp ≈ 5 μm for d = 254 μm22.

Figure 14: Variation with point to plane gap Li of the ratio A2=I/(V-Vo)2 where the continuous line indicates a L-

0.6 variation law for Li < 800 μm and the dashed one a L-1.2 variation for Li ≥ 800 μm at a low and steady liquid injection velocity of uinj ≈ 6 m/s for d = 254 μm using Diesel and the version 3 nozzle design with rp ≈ 5 μm22.

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Figure 15: Total current IT versus the applied voltage, V/Vmax for different volume flow rate at Li/d = 1.0 using Diesel and the version 3 nozzle design with rp ≈ 5 – 20 μm for (a) d = 116 μm, (b) d = 140 μm, and (c) d = 254

μm and (d) using kerosene and the version 2 nozzle design with rp ≈ 60 μm for d = 150 μm22.

Figure 16: Spray specific charge versus applied potential using kerosene: version 1 nozzle, Li/d=6.6, L0/d=2, d=500μm20.

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Figure 17: Maximum spray specific charge versus atomizer Reynolds number using kerosene: version 2 nozzle, Li/d=1.0, L0/d=2 for orifice diameters of 150μm, 250μm, 300μm and 500μm20.

Figure 18: Surface electric field strength versus atomizer Reynolds number using kerosene: version 2 nozzle, Li/d=1.0, L0/d=2 for orifice diameters of 150μm, 250μm, 300μm and 500μm20.

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Figure 19: Critical spray specific charge qV versus Re for version 3 nozzle design using Diesel no.2 with rp ≈ 5 – 20 μm for d = 254 ( ), 140 ( ) and 116 ( ) μm; from reference18 using kerosene and version 2 for d = 500 μm

with rp = 60 ( ) and 1 ( ) μm; with rp = 60 μm for 300 ( ), 250 (+) and 150 ( ) μm; using Diesel no.1 for d =

250 μm (⊗); and for version 1 for d = 500 μm using Diesel no.1 with rp = 60 (▦) and 4 (◊) μm; with rp = 60 μm using kerosene (Δ) and white spirit (˛) for d = 500 μm; and for d = 1000 μm using kerosene ( ); and for version

1 using kerosene for d = 1500 (◍), 1000 ( ) and 630 ( ) μm; from Spray Triode of reference91 using Marcol-

87 for d = 300 μm (∩), and reference30 using mineral oil for d = 422 μm ( ) and Diesel no.2 for d = 173 μm (O)22.

Figure 20: Time averaged free spray dynamics, Q=0.5mL/s, qV=-1.20C/m3, d=250μm18.

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Figure 21: Time averaged free spray dynamics, Q=0.5cmL/s, qV=-3.00C/m3, d=150μm18.

Figure 22: Time averaged spray dynamics near and earthed corner and an earthed cylinder, Q=0.5mL/s, qV=-

1.20C/m3, d=250μm18.

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Figure 23: Time averaged spray dynamics near a pair of earthed co-axial rings, Q=1.6mL/s, qV=-1.80C/m3,

d=250μm18.

Figure 24: Time averaged spray dynamics near a pair of co-axial rings, upper +7kV, lower -7kV, Q=0.5mL/s,

qV=-1.20C/m3, d=250μm18.

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Figure 25: Time averaged spray dynamics near a pair of co-axial rings, upper -7kV, lower +7kV, Q=0.5mL/s,

qV=-1.20C/m3, d=250μm18.

Figure 26: Drop diameter frequency distributions at x=0.12m, r=0.00m for (a) case 1, (b) case 2, (c) case 3, (d)

case 4 and (e) case 5, all defined in table 3.

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Figure 27: Spatial distribution of arithmetic mean diameter for (a) case 2, (b) case 4 and (c) case 5 all defined in

table 3.

Figure 28: (a) Axial and (b) Radial velocity profiles for Q=0.5mL/s, qv=-1.20C/m3 and Q=1.67mL/s, qv=-

1.80C/m3 for d=250μm.

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Figure 29: Combustion of kerosene sprays for a d=500μm atomizer with pilot flame for (a) case 1, (b) case 2

and (c) case 3, all defined in table 9 of reference78.

Figure 30: Combustion of a kerosene spray for a d=150μm atomizer (case 6 of table 9) without pilot, (a) spray

flame, (b) close up of flame seat78.

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Figure 31: Diesel oil combustion. Flow conditions as case 7 of table 9 of reference78.

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Figure 32: Effect of a positive external electric field on spray flame stability. Flow conditions as case 4 of table 7.1. (a) Vr-=+2kV, Vr+=0kV, (b) Vr-=+4kV, Vr+=0kV, (c)Vr-=+6kV, Vr+=0kV, (d) Vr-=+8kV, Vr+=0kV, (e) Vr-

=+10kV, Vr+=0kV 78.

Figure 33: Effect of a negative external electric field on spray flame stability. Flow conditions as case 4 of table 7.1. (a) Vr-=-2kV, Vr+=0kV, (b) Vr-=-4kV, Vr+=0kV, (c)Vr-=-6kV, Vr+=0kV, (d) Vr-=-8kV, Vr+=0kV, (e) Vr-=-

10kV, Vr+=0kV, (f) Vr-=-15kV, Vr+=0kV 78.

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