# Experimental Characterization of Superheated Ammonia Spray From a Single-Hole ECN Spray M Injector

Karl Oskar Bjørgen, Anthony Desclaux, Christine Mounaïm-Rousselle,
Camille Hespel

► To cite this version:

Karl Oskar Bjørgen, Anthony Desclaux, Christine Mounaïm-Rousselle, Camille Hespel. Experimental Characterization of Superheated Ammonia Spray From a Single-Hole ECN Spray M Injector. ASME 2024 ICE Forward Conference, ASME, Oct 2024, San antonio, Texas, United States. 10.1115/ICEF2024-140903. [hal-04904056]

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# EXPERIMENTAL CHARACTERIZATION OF SUPERHEATED AMMONIA SPRAY FROM A SINGLE-HOLE ECN SPRAY M INJECTOR

Karl O. P. Bjørgen<fnref n="1" />, Anthony Desclaux<fnref n="2" />, Christine Mounaïm-Rousselle<fnref n="2" />, Camille Hespel<fnref n="2" />,*

## ABSTRACT

The growing recognition of ammonia as a carbon-free alternative fuel for both transportation and power generation is underscored by numerous studies. While the use of gaseous ammonia injection has been extensively considered, research on the direct injection of liquid ammonia remains relatively scarce, especially in the case of internal combustion engines (ICE). The application of liquid ammonia in gas turbines offers economic and size-related advantages over gaseous ammonia injection. But the liquid ammonia's instantaneous spray flash-boiling and its high latent heat of vaporization induce the necessity of employing preheated swirling air to enhance flame stability, mitigating the strong cooling effect induced. The preference for direct injection of liquid ammonia is driven by its efficacy in controlling in-cylinder air-fuel ratios and optimizing thermal efficiency. This study presents macroscopic and near-field Schlieren measurements of a 1-hole ECN Spray M, representative of direct liquid injector for engine, with ammonia for different degrees of superheating. Spray angles and penetration length are provided as a function of the superheat degree with also a focus near the injector nozzle. Additionally, in-spray temperature measurements are presented, providing the first valuable information for modeling purposes.

**Keywords:** liquid ammonia, spray, flash boiling

## NOMENCLATURE

$a_0$ Droplet surface area [mm<fnref n="2" />]

$A_{S/2}$ Projected spray area at half penetration length [m<fnref n="2" />]

$D$ Nozzle hole diameter [m]

$k_m$ Boltzmann constant [J/K]

$L$ Nozzle hole length [m]

$M$ Molar mass [g/mol]

$N_a$ Avogadro's number [-]

$P_c$ Ambient gas pressure [bar]

$P_{inj}$ Injection pressure [bar]

$P_{sat}$ Saturation vapour pressure [bar]

$R$ Radial distance [m]

$R_p$ Superheat degree [-]

$S$ Spray penetration length [m]

$T_c$ Ambient gas temperature [°C]

$T_f$ Fuel temperature [°C]

$v_m$ Volume of molecule [m³]

$Z$ Axial distance [m]

$\alpha$ Spray cone angle at half penetration length [°]

$\Theta$ Dimensionless surface tension [-]

$\rho_c$ Ambient gas density [kg m⁻³]

$\rho_l$ Liquid density [kg m⁻³]

$\chi$ Non-dimensional number for the energy barrier [-]

## 1. INTRODUCTION

In the global pursuit of decarbonization, reducing greenhouse gas emissions is imperative across various sectors, including transportation and electricity generation. While strides have been made in electrifying lighter vehicles with battery electric propulsion and plug-in hybrids, challenges persist in fully transitioning heavy-duty applications such as commercial trucks, ocean marine vessels, off-road machines, and long-haul airplanes [1]. However, the urgency to address emissions extends beyond the transportation sector; it also encompasses the need for carbon-neutral or carbon-free options in dispatchable electricity production [2].

The growing recognition of ammonia as an appealing carbon-free alternative fuel for both transportation and power generation is underscored by several reviews [3, 4]. While the injection of gaseous ammonia has been first considered, research on the injection of liquid ammonia remains relatively scarce for both internal combustion engines (ICE) and gas turbines (GT) applications. In the context of ICE, the preference for direct injection of liquid ammonia is driven by its efficiency in controlling in-cylinder air-fuel ratios and optimizing thermal efficiency. Pelé et al. further highlighted the potential for achieving an optimized trade-off

<footnote n="1">¹NTNU - Department of Energy and Process Engineering, Trondheim, Norway</footnote>

<footnote n="2">²University of Orléans - PRISME, Orléans, France</footnote>

<footnote>*Corresponding author: camille.hespel@univ-orleans.fr</footnote>
between unburnt ammonia and NOx emissions through a spe-
cific direct liquid ammonia injection strategy in a single-cylinder
direct-injection spark-ignition engine [5]. The application of
liquid ammonia in gas turbines offers economic and size-related
advantages over gaseous ammonia injection. Okafor et al.'s study
emphasized the necessity of employing preheated swirling air to
enhance flame stability, mitigating the strong cooling effect in-
duced by liquid ammonia's instantaneous spray flash-boiling and
its high latent heat of vaporization [6]. The flash-boiling phe-
nomenon affects the stabilization of ammonia flame, underscor-
ing its importance in the design considerations for both internal
engines and gas turbines.

Controlling liquid ammonia injection will be done by de-
signing a pre-designed optimum thermal system. For that, CFD
models have to be improved to predict the ammonia spray vapor-
ization process well by comparing them with experimental data.
In the literature, only a few experimental studies have focused on
the characteristics of liquid ammonia spray. Recently, Li et al. [7]
investigated the injection of liquid ammonia using a single-hole
injector (280 µm nozzle diameter (D) and nozzle length-diameter
ratio (L/D) of 6). The liquid, pressurized to 300 bar and at a
temperature varying between 308 and 338 K, is injected into a
pressurized medium varying from 1.0 to 29.5 bar. Additionally,
Liu et al. [8] studied liquid ammonia injections via a single-hole
injector for a range of superheat degrees (R<sub>p</sub> = P<sub>sat</sub>/P<sub>c</sub>). They
found that going from transition flash boiling to flare flash boil-
ing, the spray radially expands, leading to large spray cone angles.
The droplet size was also found to decrease significantly when R<sub>p</sub>
was increased. Similar findings were made by Shen and Leach
[9] for liquid ammonia in a hollow cone injector. Colson et al.
[10] have also studied the injection of liquid ammonia by mod-
ifying orifice characteristics and injection conditions. Different
orifice geometries, i.e. L/D ratio, at a relatively low injection
pressure (9-12 bar) were investigated as one factor affecting the
flash boiling phenomenon as a function of (R<sub>p</sub>). Several regimes
were identified: the region without flash boiling for R<sub>p</sub> < 1, the
region of initial flash boiling for 1 < R<sub>p</sub> < 2, the transient regime
2 < R<sub>p</sub> < 5, and the complete flash boiling regime for R<sub>p</sub> > 5.
Colson et al. showed that for a geometry of the injector orifice,
L/D = 2.5, no flash boiling was observed even if the R<sub>p</sub> was
much higher than 1, whereas with an orifice of L/D = 10, the
spray was completely atomized. Therefore, this study revealed
that the transition to flashing of a liquid ammonia spray from a
single-hole nozzle could not easily be predicted using existing
criteria as summarized in [11].

In summary, as underlined by Giles et al. [12], in the case of
liquid injection of ammonia in IC engines, the injection pressure is
very high in comparison to that in gas turbines, and the 'ambient'
boundary conditions are also strongly different to the case of a
gas turbine. In IC engines, the injection can be done at elevated
pressure and/or temperature environments as a function of the
injection strategy and the injector type. The injection durations
and pressures can also vary, and even multiple injection strategies
can be considered, as in Pelé et al. [5].

There is a lack of studies characterizing liquid ammonia
sprays under these conditions, where the captured spray and va-
porization process is used for enhancing the accuracy of model-

ing, as recently highlighted by Zembi et al. [13], and addressed by
the Engine Combustion Network during workshop ECN9 [14].
Therefore, this study aims to investigate ammonia sprays un-
der superheated conditions by varying two parameters, i.e. the
ambient gas pressure and the ammonia fuel temperature. The
experiments are split into three parts, i.e., macroscopic Schlieren
measurements, near-field Schlieren measurements, and measure-
ments of the temperature inside the ammonia spray.

## 2. METHODS

2.1 Experimental setup

The test rig used was a stainless steel spherical constant
volume vessel of 4.2 L, with optical accesses via four quartz
windows, each with a diameter of 100 mm. External heaters are
installed on the vessel, keeping a constant wall temperature of
a maximum 200 °C. From the top of the vessel, an injector is
vertically mounted with an injector cooling jacket, regulating
the temperature of the injected fuel between -10 to +50 °C. The gas
system is automatically operated via a LabView control system,
where the desired operating conditions are set and executed. The
filling procedure is as follows: a vacuum pump extracts gas from
the vessel; a mass flow controller fills the vessel with a predefined
mass of synthetic air; an energizing signal is sent to the injector
and also a camera trigger; and the vessel is evacuated. The
minimum and maximum ambient gas pressure in the vessel is 0.1
bar and 10 bar, respectively.

The injector used was a single-hole gasoline direct injec-
tor with the ECN Spray M geometry (Delphi serial number:
038450BB0411213746PL). The injector is energized using an
injector driver (EFS 8427), using the recommended current pro-
file as defined by the ECN. The energizing duration was kept
constant at 7.00 ms. A summary of the thermodynamic and oper-
ating conditions can be seen in Table 1. The fuel system consists
of a high-pressure tank with liquid ammonia, pressurized by he-
lium up to 200 bar. The injection pressure is measured using a
pressure transducer installed upstream of the injector.

As the study aimed to investigate ammonia sprays under
superheated conditions, two parameters were varied, i.e., the am-
bient gas pressure and the ammonia fuel temperature. The degree
of superheating can be quantified by calculating R<sub>p</sub> for an ambient
gas pressure variation between 0.15 bar to 7 bar for the near-field
Schlieren measurements and 1 bar to 7 bar for the macroscopic
measurements. An overview of the conditions is provided in
Figure 1, where the variation in ambient pressure and fuel tem-
perature for the macroscopic Schlieren, near-field Schlieren and
temperature measurements are shown. For the macroscopic and
near-field Schlieren measurements, each condition was repeated
10 times<fnref n="1" />.

## 2.2 Optical arrangement

Two optical arrangements were used to capture the macro-
scopic and near-field characteristics of the ammonia spray, both
based on the Schlieren principle, see Figure 2. Both liquid and va-
por phases were detected in these systems. A high-power white

<footnote n="1">¹Except for the 4 bar and 7 bar conditions for the near-field, where only 5 and 3 repetitions were performed, respectively.</footnote>
<table><caption>TABLE 1: THERMODYNAMIC AND OPERATING CONDITIONS</caption><thead><tr><th>Parameter</th><th>Value</th><th>Unit</th></tr></thead><tbody><tr><td>Ambient gas</td><td>Synthetic air</td><td></td></tr><tr><td>Fuel</td><td>Ammonia</td><td></td></tr><tr><td>Ambient gas pressure (P<sub>c</sub>)</td><td>0.15-7.0</td><td>bar</td></tr><tr><td>Ambient gas temperature (T<sub>c</sub>)</td><td>28.0 ± 1.5</td><td>°C</td></tr><tr><td>Ambient gas density (ρ<sub>c</sub>)</td><td>0.11-8.00</td><td>kg/m<sup>3</sup></td></tr><tr><td>NH<sub>3</sub> fuel temperature (T<sub>f</sub>)</td><td>-10 - +50</td><td>°C</td></tr><tr><td>Injection pressure (P<sub>inj</sub>)</td><td>70</td><td>bar</td></tr><tr><td>Energizing duration</td><td>7.00</td><td>ms</td></tr></tbody></table>

<figure><img src="image_3.png" /><figcaption>FIGURE 1: EXPERIMENTAL TEST MATRIX.</figcaption></figure>

light emitting diode (LED) operated under continuous current mode was used in both optical setups. For the macroscopic measurements, a Z-type two-mirror Schlieren system was used, where a parabolic mirror (d = 107.95 mm, f = 836.6 mm) collimated the LED light and guided it through the vessel. The light was thereby decollimated via the same type of parabolic mirror as previously and guided through an adjustable pinhole for light cut-off in all directions. The light was lastly focused onto the camera sensor using a lens with a focal length of 200 mm. For the near-field measurements, a dual-field-lens system was used, where the LED light was collimated using an 82 mm diameter lens (f = 300 mm) and decollimated using the same type of lens on the opposite side of the vessel. A horizontally mounted knife's edge was placed in the focal point for light cut-off, making the system sensitive density gradients in the vertical direction (along the spray axis). The light is lastly focused onto the camera sensor using a parabolic mirror (d = 107.95 mm, f = 836.6 mm). The near-field Schlieren system resulted in a magnification of 3.6, meaning that the injector nozzle diameter of 0.170 mm is projected to 0.612 mm on the image sensor, equating to 7.758 µm/pixel. The camera settings used for macroscopic and near-field measurements are given in Table 2.

## 2.3 Image processing

The spray geometry of the macroscopic Schlieren images was determined by using an in-house MATLAB image processing

<figure><img src="image_4.png" /><figcaption>Near-field Schlieren</figcaption></figure>

<figure><img src="image_5.png" /><figcaption>FIGURE 2: THE OPTICAL SETUPS FOR THE MACROSCOPIC AND NEAR-FIELD SCHLIEREN.</figcaption></figure>

<table><caption>TABLE 2: CAMERA SETTINGS FOR SCHLIEREN MEASUREMENTS.</caption><thead><tr><th>Parameter</th><th>Macroscopic</th><th>Near-field</th></tr></thead><tbody><tr><td>Camera model</td><td>Phantom V1611</td><td>Phantom V2512</td></tr><tr><td>Frame rate</td><td>45000 fps</td><td>51000 fps</td></tr><tr><td>Resolution</td><td>256x800</td><td>768x576</td></tr><tr><td>Exposure time</td><td>5 µs</td><td>10 µs</td></tr><tr><td>Physical resolution</td><td>106.053 µm/px</td><td>7.758 µm/px</td></tr></tbody></table>

program. In this process, the raw images underwent binarization to extract essential information regarding spray development, including spray penetration and spray angle. Initially, thirteen background images with no spray were utilized to compute an average background, which was then subtracted to isolate the spray. Following background subtraction, the Otsu thresholding method [15] was applied for binarization, effectively identifying the liquid phase of the spray and facilitating the determination of the mentioned parameters. The spray cone angle (α) is defined as the angle from the nozzle outlet to half the spray's penetration and is calculated from

$$ \alpha = 2 \arctan \frac{A_{S/2}}{(S/2)^2}, \quad (1) $$

where S is the spray penetration, quantified as the vertical distance between the nozzle exit and the furthest point reached by the liquid spray, and A<sub>S/2</sub> the projected spray area until the half penetration length.
<figure><img src="image_6.png" /><figcaption>FIGURE 3: POST-PROCESSING ROUTINE OF THE NEAR-FIELD SCHLIEREN MEASUREMENT FOR y = 2D AND 10D. LEFT FIGURE: ENSEMBLE-AVERAGED AND BACKGROUND SUBTRACTED IMAGE OF THE SPRAY; CENTER FIGURE: EXTRACTED NEGATIVE VALUES; RIGHT FIGURE: EXTRACTED POSITIVE VALUES FROM MEASUREMENT IMAGE.</figcaption></figure>

<figure><img src="image_7.png" /><figcaption>FIGURE 4: DOWNSTREAM POSITIONS FOR IN-SPRAY TEMPERATURE MEASUREMENTS.</figcaption></figure>

For the near-field Schlieren measurements, the structuring and postprocessing of the data were performed using Python. The background image with no spray was computed using the first fifteen images for each measurement. Since the near-field Schlieren measurements resulted in the vapor phase appearing as positive values and the liquid phase as negative values in the background subtracted images, these could be separated by extracting the positive and negative values and evaluating their spray angles separately, which can be seen in Figure 3. In order to detect the spray boundary of the vapor phase, Yen's thresholding method [16] was used, while for the liquid phase, a fixed value was chosen, i.e. 90 % of the difference between the background brightness level and the dark level of the camera. The spray angle was calculated as the angle between the radial location of the left and right threshold lines for y locations y = 2D and y = 10D, where D = 170 µm is the small diameter of the nozzle hole. The spray angle is calculated individually for all images taken between 1.96 ms and 6.86 ms after the start of the energizing signal, representing the steady state period during the injection.

Thereby, the mean and standard deviation for all images and repetitions are calculated.

## 2.4 In-spray temperature measurements

For measuring the in-spray temperature during injection, three type K thermocouples (thickness of 25 µm) were set in the vessel, see Figure 4. Ten operating conditions were chosen in order to study the effect of superheat degree on the spray temperature, i.e. for $0.6 < R_p < 15.0$. The energizing duration during the in-spray temperature measurement was 7.0 ms, ensuring that the thermocouples were able to capture the correct temperature and not being affected by the thermal inertia of the thermocouple. The thermocouples were mounted on a translation stage so that the radial position of the thermocouple measurement could be varied. For the $R_p = 15.0$, a finer radial measurement grid was used in order to obtain more details on the radial temperature gradient. The temperature was measured at three downstream axial positions, i.e. $Z = 9$ mm, 32 mm and 55 mm, where $Z = 0$ mm corresponds to the nozzle tip location. The accuracy of thermocouple measurement was estimated to be ±1.5 °C.

# 3. RESULTS

## 3.1 Macroscopic spray characteristics

Figure 5 shows the temporal development of the spray penetration length for 2 bar ambient pressure conditions with varying fuel temperatures, corresponding to superheat degrees between 1.53 and 9.64. For all conditions, the spray penetration develops similarly until 0.2 ms after the start of injection. Thereafter, the spray penetration develops slower when fuel temperature increases, i.e., for increased superheat degree ($R_p$). This could be directly linked with the increasing tendency of flash boiling. In Figure 6, the fuel temperature is fixed at -10 °C while the ambient pressure is varied from 1 bar to 5 bar, corresponding to ambient densities between 1.1 and 5.8 kg/m³. In this case, the spray penetration is mostly affected by the increasing ambient density, causing the spray to slow down. However, for the most flashing case, i.e. 1 bar ambient pressure ($R_p$ = 3.01), the spray penetration is shorter than that of the 2 bar ambient pressure case ($R_p$ = 1.53). This indicates that the superheat degree has
<figure><img src="image_8.png" /><figcaption>FIGURE 5: THE SPRAY PENETRATION LENGTH AS A FUNCTION OF TIME FOR 2 BAR CHAMBER PRESSURE WITH VARYING FUEL TEMPERATURE.</figcaption></figure>

<figure><img src="image_9.png" /><figcaption>FIGURE 6: THE SPRAY PENETRATION LENGTH AS A FUNCTION OF TIME FOR -10 °C FUEL TEMPERATURE WITH VARYING CHAMBER PRESSURE.</figcaption></figure>

a greater impact on the spray penetration than the air density for $R_p > 3$.

The effect of ambient density and $R_p$ on spray penetration and spray angle is shown in Figure 7 and 8, respectively. For low $R_p$, an increased ambient density clearly shortens the penetration, while for intermediate values of $R_p$, a smaller influence is observed. For higher $R_p$, the spray penetration stabilizes. The variation in the influence of ambient density on spray penetration deviates from non-flashing sprays, suggesting that the degree of superheat has an impact on the spray penetration. For the measured spray angles for low degrees of superheat ($R_p < 3$), there is a clear influence of ambient density, where wider angles occur for higher ambient densities. As $R_p$ increases, the spray angle increases too, but the ambient density effect diminishes. For $R_p > 8$, the spray angle stabilizes due to the increasing effect of flash boiling.

<figure><img src="image_10.png" /><figcaption>FIGURE 7: THE SPRAY PENETRATION LENGTH AT $t = 0.8$ SECONDS AS A FUNCTION OF SUPERHEAT DEGREE ($R_p$).</figcaption></figure>

<figure><img src="image_11.png" /><figcaption>FIGURE 8: THE SPRAY ANGLE AT HALF PENETRATION LENGTH AS A FUNCTION OF SUPERHEAT DEGREE ($R_p$).</figcaption></figure>

Figure 9 plots the calculated time exponents of the initial injection period. The power curve $S(t) = at^n$ is fitted to the temporal penetration measurement between $t = 0$ ms and $t = 0.3$ ms by the least-square method, where $S$ is the spray penetration, $a$ is a constant, and $n$ is the time exponent. During the initial injection period, the spray penetration is not only governed by the spray momentum at the nozzle exit but also by the nucleation of bubbles near the nozzle exit. For higher $R_p$, the axial spray momentum decreases, resulting in the time exponent to stabilize around 0.45 after $R_p = 10$. For $R_p$ higher than 3, there is a significant effect of the ambient density on the time exponent, which is different from what is observed for the spray angle.

Comparing the outcomes of spray penetration and the fitted time exponent with those observed for the spray angle, it becomes evident that the reduction in axial spray penetration correlates with the increase in lateral spray expansion. This expansion
<figure><img src="image_12.png" /><figcaption>FIGURE 9: THE TIME EXPONENT AS A FUNCTION OF SUPERHEAT DEGREE ($R_p$).</figcaption></figure>

<figure><img src="image_13.png" /><figcaption>FIGURE 10: SPRAY ANGLE AS FUNCTION OF $R_p$ FOR $y = 2D$. THE VAPOR AND LIQUID PHASES ARE PLOTTED TOGETHER.</figcaption></figure>

primarily relies on superheat degrees of $R_p > 4$.

These findings confirm the subdivision of the flashing region of superheated ammonia into three distinct regimes: initial flashing, transitional flashing, and flare flashing regions. However, the observed impact of air density shows that not only $R_p$ should be considered for macroscopic spray characteristics under flashing conditions.

## 3.2 Near-field spray characteristics

The near-field spray measurements are presented in Figure 13 for all conditions. The ammonia spray is widened by increasing fuel temperatures and decreasing chamber pressures. For some conditions, for e.g. $P_c, T_f = (4 \text{ bar}, 20 \°C)$, there is no widening of the spray, indicating little or no flash boiling or mechanical break up of the injected fuel. On the contrary, e.g. $P_c, T_f = (0.3 \text{ bar}, 30 \°C)$, the spray has a very sudden widening of the spray as the fuel exits the nozzle. The effects of the two mechanisms

<figure><img src="image_14.png" /><figcaption>FIGURE 11: SPRAY ANGLE AS FUNCTION OF $R_p$ FOR $y = 10D$ FOR THE LIQUID PHASE.</figcaption></figure>

<figure><img src="image_15.png" /><figcaption>FIGURE 12: THE SUPERHEAT DEGREE AS FUNCTION OF $\chi$. THE LIMITS AT $\chi = 1$ AND $\chi = 4$ INDICATE THE TRANSITION REGIME ([17]) ARE SHOWN.</figcaption></figure>

are indistinguishable, requiring an analysis of the spray angle and the thermodynamic conditions. For non-flashing conditions, the liquid fuel breaks up when colliding with air, following the well-known jet break-up regimes depending on the injection pressure, velocity and chamber conditions, while under flash boiling conditions, internal bubbles form inside individual droplets, causing them to break up further, leading to additional widening of the spray.

In order to detect the transition from non-flashing to fully flashing conditions, the spray angles for all tested conditions are plotted as a function of $R_p$ at the downstream location $y = 2D$ and $y = 10D$, as shown in Figure 10 and 11. Note that both the vapor and liquid phase spray angles are presented in Figure 10. The measured vapor phase spray angle is always larger than the liquid spray angle for similar $R_p$. The liquid spray angle increases monotonically for $R_p$ higher than approximately 2, up to $R_p = 10$, thereafter, it levels off for increasing $R_p$ up to 40. A
<figure><img src="image_16.png" /><figcaption>FIGURE 13: AVERAGED SPRAY IMAGES FOR ALL FUEL TEMPERATURES AND CHAMBER PRESSURE MEASURED USING THE NEAR-FIELD SCHLIEREN SETUP. THE IMAGES MARKED WITH A RED RECTANGLE HAVE SIMILAR SUPERHEAT DEGREES ($R_p = 8-10$).</figcaption></figure>

similar trend is observed for the vapor spray angle, although for $R_p$ lower than 5, it was not possible to distinguish between the vapor and liquid phases. Comparing these measurements (liquid phase for $y = 2D$ and $y = 10D$) to the macroscopic measurements in Figure 8, there is no influence of ambient density for the low $R_p$ conditions, which can be explained by the relatively small time scale for the near-field measurements. This means that the radial expansion is mostly attributed to the flash boiling phenomenon.

The flash boiling phenomenon is hypothesized to be governed by Classical Nucleation Theory (CNT) via the rate of bubble nucleation inside a droplet in the spray. Following Girshick and Chiu [18] and [17] this can be represented by the non-dimensional number $\chi$ as

$$ \chi = \frac{4\Theta^3}{27(lnR_p)^2} \quad (2) $$

$\Theta$ is the dimensionless surface tension defined as $\Theta = (a_0\sigma)/(k_bT_f)$, where $a_0$ is the surface area ($a_0 = (36\pi)^{1/3}v_m^{2/3}$), $k_b$ is the Boltzmann constant, $v_m$ is the volume of a molecule

approximated as $v_m = M/(\rho_l N_\alpha)$, with M denoting the molar mass of ammonia, $\rho_l$ is the liquid density of ammonia, and $N_\alpha$ is the Avogadro Constant. The non-dimensional number $\chi$ is small when the chemical potential for phase transition is large compared to the work required to form a new surface, indicating the onset of vapor nucleation and, hence, the onset of flashing conditions. From Lamanna et al. [17], the onset for flashing conditions are found for $1 \le \chi \le 4$, where all measurements with $\chi < 1$ was identified as fully flashing, independent of fluids tested.

Figure 12 shows the relation between $\chi$ and $R_p$, where the threshold for the onset of flash boiling conditions ($\chi = 4$) and for fully flashing conditions ($\chi = 1$) are displayed. For $R_p \le 2.5$ all conditions fall under the non-flashing regime. For values of $7 \le R_p < 20$, the tested conditions are distributed across the transition region and the fully flashing regime. For values of $R_p \ge 20$, all conditions fall under the fully flashing regime. An interesting point is $P_c, T_f = (4 \text{ bar}, 50 \text{ $^\circ$C})$, where $R_p = 4.92$ and $\chi = 0.61$. Based on Figure 10, the low $R_p$ suggests that the condition is in a transitional flashing regime, while the $\chi$ is below
<figure><img src="image_17.png" /><figcaption>FIGURE 14: SHOCK WAVES OBSERVED FOR HIGH R<sub>p</sub> CASES.</figcaption></figure>

<figure><img src="image_18.png" /><figcaption>FIGURE 15: MINIMUM TEMPERATURE CONTOUR PLOT OVERLAYED WITH THE MACROSCOPIC SPRAY SHAPE. T<sub>m</sub> INDICATES THE LOWEST MINIMUM TEMPERATURE ALONG Z = 55 MM.</figcaption></figure>

unity, suggesting that the spray is fully flashing. The degree of flashing for these points (R<sub>p</sub> < 7 and χ < 1) are therefore difficult to determine based on this framework.

Looking into the measurements with 7 < R<sub>p</sub> < 10, i.e. P<sub>c</sub>, T<sub>f</sub> = (2 bar, 40 °C), (1 bar, 20 °C) and (0.5 bar, 0 °C), (0.3 bar, -10 °C) and (2 bar, 50 °C) (indicated with red rectangles in Figure 13), they are distributed from χ = 0.3 to χ = 3 in Figure 12. This means that for a similar degree of superheating, the sprays fall under different flashing regimes according to χ [17]. The spray angles in Figure 10 for these points are similar, and the different degrees of flash boiling are therefore not observed using this measure. By analyzing the spray morphology for these conditions, the spray shape differs significantly even for similar R<sub>p</sub> and with similar spray angles. For the high χ cases (e.g. P<sub>c</sub>, T<sub>f</sub> = (0.3 bar, -10 °C)), there is a less dense liquid core compared to the lower χ cases, suggesting that the number of droplets is lower and/or the droplet sizes are smaller, which could indicate a higher degree of flash boiling or evaporation. However, according to Lamanna et al. [17], a higher value of χ leads to a

less flash boiling case, which is the opposite of what is observed
in this case.

3.3 Shocks in the ammonia spray

For very high degrees of superheat for the highest fuel tem-
perature condition, shock waves were observed using the near-
field Schlieren setup. Figure 14 shows near-field Schlieren mea-
surement for R<sub>p</sub> = 64.6, where weak shocks can be seen ap-
proximately 1 mm downstream of the nozzle exit. As the R<sub>p</sub>
increases by decreasing the chamber pressure, the spray becomes
wider, and the liquid core becomes fainter, enabling a clearer
observation of the shock waves. Comparing the location of the
shocks for different chamber pressures, higher pressures increase
the distance from the nozzle exit to the shock wave. Additionally,
the width of the shock structure is observed to broaden as R<sub>p</sub>
increases, which is in line with previous findings by Zhang et
al. [19] for n-hexane sprays. Although shock waves were not
observed for cases with R<sub>p</sub> lower than 64.6, there could still be
shocks closer to the nozzle exit hidden in the liquid core.
## 3.4 In-spray temperature

Figure 15 shows a contour plot of the in-spray temperature measurements overlayed with the spray boundary based on the macroscopic Schlieren measurement. The contour plot is based on the minimum temperature during a 7 ms duration injection, measured for three downstream positions and 4-6 radial positions<fnref n="2" />. The decrease in temperature observed in the spray core is caused by the evaporating ammonia, where the cooling is especially strong due to the very high latent heat of vaporization of ammonia. The in-spray temperature measurements follow the liquid spray contour closely, where a sudden decrease in temperature is observed within the spray boundary. The lowest temperature measured was -66 – -64 °C for the 1 bar ambient pressure cases, and the highest was -41 °C for the 7 bar pressure case.

The in-spray temperature was found to be dependent on the ambient pressure in the chamber, where increasing pressure increased the in-spray temperature. This was also observed by Somarathne et al. [20] and explained by the increase in ammonia saturation temperature for increased pressure. Hence, increased ambient pressure can reduce the need for heat addition, resulting in improved spray combustion characteristics.

Comparing the cases $P_c, T_f$ = (1 bar, 0 °C) and (3 bar, 40 °C), with $R_p$ = 4.3 and 5.3, respectively, the lowest temperature measured was found for the condition with the lower $R_p$.

For the conditions with the highest $R_p$, i.e. 5.3, 6.2, 8.8 and 15.0, it is observed that the sudden drop in temperature coincides well with the spray boundary for the upstream positions $Z = 9$ mm and $Z = 32$ mm, while for $Z = 55$ mm, the temperature drop occurs closer to the spray center axis. This could be explained by the wider spray angle for these conditions, causing more mixing with the ambient air and thereby increasing the in-spray temperature close to the radial edge of the spray.

## 4. CONCLUSION

In this study, detailed measurements of a single spray with liquid ammonia injected under superheated conditions are presented for a single-hole injector referenced by ECN. The results show that on a macroscopic scale, the spray morphology is significantly affected by ammonia's degree of superheat. For $R_p < 3$, the spray angle depends mainly on the ambient density, where higher densities lead to larger angles. For $R_p > 4$, the radial expansion of the spray increases monotonically and is not affected by changes in ambient air density. It is also observed that as the spray widens for higher $R_p$, the spray penetration shortens, suggesting that these effects are correlated. The near-field measurement of the spray angle was found to have a similar dependency on the $R_p$ for high values of $R_p$, although, for low $R_p$, no influence of ambient density was observed. This indicates that the ambient density only influences the spray widening in the far field for low $R_p$ conditions.

In the near-field measurements, conditions with $7 < R_p < 10$ resulted in similar spray angles. However, when studying the spray structure, the liquid core appeared much less dense for the low ambient pressure cases compared to the high ambient pressure cases, which could indicate that there is a parameter

leading to flash boiling which is not captured by $R_p$ or the spray angle.

In-spray temperature measurements revealed that the minimum temperature reached during the steady state in the spray core was correlated with the ambient air pressure, where lower pressure led to lower in-spray temperatures, while higher superheat conditions did not necessarily lead to lower temperatures. Also, for high $R_p$ cases, the radial temperature drop occurs closer to the spray axis compared to low $R_p$ cases, which could be due to the mixing of ambient air along the spray periphery for the widened, flash boiling sprays.

Shocks were observed in the ammonia spray for cases with $R_p > 65$. For increasing $R_p$, the shock structure grew in width and axial length.

## ACKNOWLEDGMENTS

The work has been partially financed by ANR for the ADONIS Project ANR-22-HTCE-0001-02 call EIG CONCERT-Japan 2021 and EEA and Norway Grants via ACTIVATE [NOR/POLNOR/ACTIVATE/0046/2019-00].

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