**4.2 Thermodynamic characteristic of DBD plasma actuations under frozen cold conditions with significant heat convection pertinent to aircraft anti**−**/de-icing**

As described above, while a number of investigations have been conducted to characterize the thermal effects of DBD plasma discharges [30–32, 40], almost all the previous studies were conducted in quiescent air at room temperature without considering frozen cold conditions and significant convective heat transfer pertinent to aircraft icing phenomena. With the experimental set up described above, a comprehensive investigation was conducted to characterize the thermodynamic characteristics of DBD plasma actuations as a function of relevant controlling parameters (e.g., applied voltage, frequency, and power input, etc...) under frozen-cold test conditions coupled with significant convective heat transfer pertinent to aircraft anti−/de-icing.

**Figure 6** shows one example of the experimental results to reveal the time evolution of the measured temperature distribution over the airfoil surface protected by the DBD plasma actuator a dry test condition (i.e., without turning on the water spray system of the ISU-IRT). For the experiment results, while the incoming airflow was set at of *U*∞ = 40 m/s and *T*∞ = −5.0°C, the AC-DBD plasma actuators were supplied by AC voltage of *V*p-p = 12.5 kV and *f* = 10 kHz, with the corresponding applied power density of *P*d = 7.8. kW/m2 . It is clearly seen that, after the plasma actuator was switched on, the temperatures over the airfoil surface were found to increase rapidly, with the local surface temperatures at the edges of the exposed electrodes being raised from −5°C to more than 25°C in less than 5 seconds. As shown clearly in **Figure 6(a)**, the surface heating was first initiated at the edges of the exposed electrodes with evident local temperature peaks (i.e., as indicated by the white strips in the temperature map over the airfoil surface). As the time goes by, more and more thermal energy was generated during the plasma discharges, as seen from the measurement results shown in **Figure 6(b)**–**(d)**. It should be noted that the maximum temperatures were always found to be located at the edges of the exposed electrodes, which agrees with the findings reported in the previous studies [40]. Meanwhile, the temperature over the exposed electrodes (i.e., copper tap) appeared to be much higher than that over the dielectric layer (i.e., Kapton film). The temperature differences between the electrode surfaces and the surface of the dielectric layer were believed to be caused by the significant difference in the thermal conductivity between the copper tape and the Kapton film (i.e., 385.0 W/m·K for copper tape *vs.* 1.57 W/m·K for Kapton film). The measured surface temperature was found to be relatively low near the airfoil leading edge in general, and increased gradually at further downstream locations, which was correlated well with the chordwise development of the convective heat transfer over the airfoil surface (i.e., the heat convection would be maximum at the airfoil leading edge, and decrease gradually in the downstream [49]).

Based on the measured temperature distributions given in **Figure 6**, the spanwise-averaged temperature profiles along the airfoil chord can be extracted, and the extracted results are given in **Figure 7**. It can be seen clearly that the spanwiseaveraged temperature profiles at the different time instances have a very similar distribution pattern, i.e., the surface temperatures were always found to reach the local peak values at the edges of the exposed electrodes, and then decrease gradually

### **Figure 6.**

*Measured temperature distribution over the airfoil surface with the DBD plasma actuator operating under a dry test condition of U∞ = 40 m/s and T∞ =* −*5°C.*

*An Experimental Investigation on the Thermodynamic Characteristics of DBD Plasma… DOI: http://dx.doi.org/10.5772/intechopen.100100*

### **Figure 7.**

*Spanwise-averaged temperature profiles along the airfoil chord with the DBD plasma actuators operating under a dry test condition of U∞ = 40 m/s and T∞ =* −*5°C.*

with the increasing distance away from the edges of the exposed electrodes. For example, at the time instance of *t* = 0.4 s, while the surface temperatures over the dielectric layer were found to have almost no change (i.e., still being frozen cold at *T*surface = −5°C), the local temperature peaks at the electrode edges were found to increase rapidly and become higher than +5.0°C. Along with the rapid temperature rise at the edges of the exposed electrodes, the temperatures over the surfaces of copper-based electrodes were also found to become much higher in comparison to those over the dielectric layer. It was suggested that the predominant heating mechanism in plasma discharges is due to the heat transfer from the plasma to the gas, which then heats up the surface of the plasma actuator through forced convection [32]. With the plasma actuator embedded over the airfoil surface was exposed in the frozen cold airflow coupled with significant convective heat transfer, the hot air originally heated by plasma discharges would not only be in contact with the dielectric layer, but also convect over the copper-based exposed electrodes. It caused the temperature rise over the surfaces of the exposed electrodes. Due to the much higher thermal conductivity of the copper-based exposed electrodes, the electrode surfaces were found to have a much faster thermal response (*i.e.*, rapid temperature increases) in comparison to that of the Kapton-based dielectric layer, as shown quantitatively from the measured temperature profiles given in **Figure 6**. As the time goes on, the temperatures over the airfoil surface were found to increase rapidly, with the maximum temperature raised to more than 35°C at 6.4 seconds after turning on the DBD plasma actuator. It can also be seen clearly that, the surface temperature around the third exposed electrode was found to be always higher than those at other locations, which is believed to be a result of the development of the thermal boundary layer over the airfoil surface, i.e., due to the effects of the significant convective hear transfer over the airfoil surface [50].

## **4.3 Comparison of the anti**−**/de-icing performance of the DBD plasma-based approach against the convention electrical surface heating methods**

With the experimental setup given in **Figure 5**, a comprehensive experimental campaign was conducted to provide a side-by-side comparison of the DBD

plasma-based approach against conventional electrical heating methods in preventing the ice formation and accretion over the airfoil surface. In performing the ice accretion experiments, ISU-IRT was operated at a prescribed frozen-cold temperature level (e.g., *T*∞ = −5°C for the present study) for at least 30 minutes in order to ensure ISU-IRT reaching a thermal steady state. Then, the DBD plasma actuator and the electrical film heater embedded over the airfoil/wing surface were switched on simultaneously for about 60 seconds to achieve a thermal equilibrium state before turning on the water spray system of ISU-IRT. After the water spray system was switched on at *t* = *t*0, the super-cooled water droplets carried by the incoming airflow would impinge onto the surface of the airfoil/wing model to start the ice accretion process. During the experiments, the high-speed imaging system and IR thermal imaging system were synchronized to record the dynamic ice accretion or anti−/de-icing process and map the corresponding surface temperature distributions over the ice accreting airfoil/wing model simultaneously.

**Figure 8** shows the typical snapshots of the dynamic ice accretion process over the airfoil surface with the same electric power supplied to the DBD plasma actuator and the electrical film heater (i.e., *P*d = 7.8 kW/m2 ) for the anti−/de-icing operation. The box in red dashed lines in the acquired images indicates the measurement window of the IR thermal imaging system. Since very similar features were observed for all the test cases, only the measurement results obtained under the test conditions of *U*∞ = 40 m/s, *T*∞ = −5°C and *LWC* = 1.0 g/m3 were shown and analyzed here for conciseness. As shown clearly in **Figure 8(a)**, right after starting the ice accretion experiments (*i.e.*, *t* = 10.0 s), since both the DBD plasma actuator and the electrical film heater had already been switched on for a while to make the surface temperatures of the airfoil/wing model being well above the freezing point of water, the supercooled water droplets were found to be heated up rapidly, upon impacting onto the heated airfoil surface. Therefore, the front surface of the airfoil model protected by the plasma actuator and the electrical film heater (*i.e.*, from the leading edge to ~30% chord length) was found to be totally ice free, with evident water runback flow observed over the airfoil surface. Driven by the boundary layer airflow over the airfoil surface, the unfrozen water was found to run back in the form of film/rivulet flows. The runback water over the airfoil surface on the electrical film heater side was found to be refrozen into ice eventually to form rivuletshaped ice structures at the downstream region of X/C ≈ 60%. In comparison,

### **Figure 8.**

*Acquired images to reveal the dynamic ice accretion process over the airfoil surface with the electric power supplied to the plasma actuator and the film heater being Pd = 7.8 kW/m<sup>2</sup> .*

### *An Experimental Investigation on the Thermodynamic Characteristics of DBD Plasma… DOI: http://dx.doi.org/10.5772/intechopen.100100*

much less runback water was observed over the airfoil surface on the DBD plasma actutor side with much fewer ice structures formed at the downstream locations.

The experimental observation suggests that, with the same electric powere inputs, the DBD plasam actuator seems to have a better anti−/de-icing performance in comparison to the conventioneal electric film heater. This can be explained by the facts that, since the airflow over the region covered by the DBD plasma actuator would be sufficiently heated due to the gas heating effects in the plasma acuatuation, a portion of the airbrone water droplets would be warming up rapidly and even evaporated as they flying through the plasam region before impacting onto with airfloil surface, resulting in the less water mass collected over the airfoil surface proptected by the DBD plasma actutor. However, with the supercooled water droplets impinging onto the airfoil surface protected by conventional electrical film heater, the thermal energy was mainly transferred from the heating elements to the airfoil surface via heat conduction. Since the input power for this cases was not sufficient to instantly evaporate the impacted water droplets (i.e., *P*d = 7.8 kW/m2 ), the impacted water droplets were found to coalesce quickly on the airfoil surface to form rivulets/film flows to transport the impacted water mass to further downstream locations, as driven by the airflow over the airfoil surface. Due to the intense convective heat transfer between the surface water and the frozen cold incoming airflow over the airfoil surface, the runback water was found to be refrozen into ice eventually, result in the formation of rivulet-shaped ice strcutures at further downstream locations.

As the time goes on, more and more super-cooled water droplets would impinge onto the airfoil surface. As a result, more water mass was found to be collected over the airfoil surface to cause the formation of more rivulet-shaped runback ice accreted at the downstream locations of the airfoil surface. As shown in **Figure 8(b)** and **(c)**, while electric power supplied to the two systems were set to be the same value of *P*d = 7.8 kW/m<sup>2</sup> , the ice structures accreted over the airfoil surface on the plasma actuator side were always found to be less than those on the electric film heater side.

In order to achieve a better anti−/de-icing performance, the electric power supplied to the plasma actuator and the electrical film heater were increased by a factor of two (*i.e.*, *P*d = 15.6 kW/m2 ) to generate more thermal energy for the anti−/de-icing. The typical snapshots of the dynamic water runback/ice accretion process with elevated power input are shown in **Figure 9**. It can be clearly seen that, with the higher power input of *P*d = 15.6 kW/m2 , the airfoil surface was found to become completely free of ice on both sides of the airfoil surface. It can also be seen that, similar to that observed for the case with relatively lower power input described above, the DBD plasma side of the airfoil surface appeared to have much less water runback in comparison to that on the electric film heater side. The rapid evaporation of the airborne water droplets as flying through the plasma region described above is believed to be the reason to cause the much less water mass collected on the plasma actuator side of the airfoil surface.

The corresponding IR thermal imaging results can reveal more details on the different working mechanisms of the two system for the anti−/de-icing operation. While **Figure 10** shows the time evolution of the measured temperature distributions over the airfoil surface before and after starting the ice accretion process, **Figure 11** gives the corresponding surface temperatures at different chordwise locations (*i.e.*, locations of *A, B, C* and *D* at *X/C* = 2.0%, 10% 18% and 45% chord respectively, as indicated in **Figure 10**) as a function of the time on the two sides of the airfoil surface (*i.e.*, plasma actuator side *vs.* electric film heater side). It can be seen clearly that, after the DBD plasma actuator was switched on for 10 seconds (i.e., at *t* = 10s), the temperatures over the exposed electrodes of the plasma

### **Figure 9.**

*Acquired images to reveal the dynamic ice accretion process over the airfoil surface with the electric power supplied to the plasma actuator and the film heater being Pd = 15.6 kW/m2 .*

### **Figure 10.**

*Time evolution of the measured temperature distributions over the airfoil surface before and after starting the ice accretion process.*

actuator were found to increase to about 10°C, while the temperatures over the dielectric surface (*i.e.*, in the spacings between the electrodes) were still quite low (i.e., below the freezing point of water), which has been discussed in the previous section. Since the electrical film heater was also switched on simultaneously, a

*An Experimental Investigation on the Thermodynamic Characteristics of DBD Plasma… DOI: http://dx.doi.org/10.5772/intechopen.100100*

**Figure 11.** *Measured surface temperatures at different chordwise locations on the airfoil surface before and after starting the ice accretion process.*

strip-patterned temperature distribution was found over the airfoil surface on electric film heater side at *t* = 10 seconds. Such a strip-like temperature distribution was due to the configuration of the etched foil resistance element encapsulated between the Polyimide films in the electric film heater.

As the time goes by, more and more thermal energy would be generated on both the plasma actuator side and the electric film heater side of the airfoil surface. At about 50 second after turning on the plasma actuator and the electric film heater, a thermal equilibrium state was found to achieve on both sides of the airfoil surface, as indicated by the flattening surface temperature profiles shown in **Figure 11(a)** and **(b)**. It can be seen clearly that, the surface temperatures over the electrical film heater were much higher than those over the DBD plasma actuator as they reached the thermal equilibrium state. The temperatures at the downstream location "C" (*i.e.*, X/D = 18%) were found to be the maximum on both sides with the measured values becoming 20°C and 90°C on the plasma actuator side and the electric film heater side, respectively. As described above, while the thermal energy generated by the electrical film heater is mainly at the heater surface through resistive heating, the primary heating mechanism in DBD plasma actuation is through gas heating and then heating up the dielectric/electrodes surfaces through direct injection, convection and radiation [32]. Therefore, with the same power input, the measured surface temperatures on the electric film heater surface were found to be much higher than those over the DBD plasma actuator. It was also found that, after the thermal equilibrium state was achieved, the temperature was higher at the locations further away from the airfoil leading edge (i.e., X/C ≈ 18%) as clearly shown in **Figures 10** and **11**. Existence of such a temperature gradient over the airfoil surface was believed to be caused by the development of the thermal boundary layer over the airfoil surface, i.e., due to the significant convective heat transfer over the airfoil surface with the maximum heat convection locating at the airfoil leading edge and decreasing gradually in the downstream region.

As indicated by the dashed line in **Figure 11**, the water spray system of ISU was switched on at *t* = 60 s to start the ice accretion process, i.e., at 60 seconds after turning on the plasma actuator and the electrical film heater. This time instant was also defined as *t*0, as given in **Figure 10(b)**. It was found that, after the super-cooled water droplets impinged onto the airfoil surface, for the time instance at *t* = *t*0 + 25 s, while the surface temperature on the electric film heater was found to decrease significantly, the temperature on the surface of the plasma actuator only dropped

slightly as shown in **Figure 10**. As the time goes on, more and more impinged water would be collected on the airfoil surface. Since the power input to the plasma actuator and the electric film heater were sufficiently high to prevent ice accretion over the airfoil surfaces (i.e., *P*d = 15.6 kW/m2 ), the mass transport and energy transfer on both sides of the airfoil surface were found to reach an equilibrium state, as indicated by the almost unchanged temperature distributions. More quantitatively, after the water droplets impinged on the airfoil surface, while the temperature drop on the airfoil surface protected by the plasma actuator was about 33% (*i.e.*, the temperature dropped from 12.0°C to 8.0°C at the location A and B, and dropped from 18.0°C to 12.0°C at the location C), the corresponding temperature decrease son the airfoil surface protected by the electrical film heater were found to be around 70% (*i.e.*, the temperature decreased from 25.0°C to 6.0°C at the location A, from 60.0°C to 20.0°C at the location B, and from 90.0°C to 25.0°C at the location C). Such significant differences in the surface temperature changes before and after the impingement of the super-cooled water droplets can be explained by the different heating mechanisms discussed in the previous section. For the electrical film heater case, since the thermal energy was mainly generated at the heater surface, and then transferred into the impinged supercooled water droplets, the measured surface temperature, therefore, appeared to drop significantly due to the great temperature differences between the heater surface and the impinged water droplets. However, for the case with the water droplets impinging onto the surface of the DBD plasma actuator, the water droplets had already been effectively warmed up through the forced heat convection as they were flying through the hot air above the plasma actuator. Since the temperatures of the water droplets would be increased substantially before impacting onto the dielectric/electrodes surface, it results in the much smaller surface temperature drops upon the impacting of the water droplets onto the airfoil surface protected by the plasma actuator, as revealed quantitatively in **Figure 11(b)**.
