IR-Thermography for Dynamic Detection of Laminar-Turbulent...

18th International Symposium on the Application of Laser and Imaging Techniques to Fluid Mechanics・LISBON | PORTUGAL ・JULY 4 – 7, 2016

IR-Thermography for Dynamic Detection of Laminar-Turbulent Transition

Bernhard Simon1,*, Adrian Filius1, Cameron Tropea1, Sven Grundmann2 1: Center of Smart Interfaces, Technische Universität Darmstadt, Germany

2: Dept. of Fluid Mechanics, University of Rostock, Germany * Correspondent author: [email protected]

Keywords: IR thermography, laminar-turbulent transition

ABSTRACT

This work investigates the potential of infrared (IR) thermography for the dynamic detection of laminar-turbulent transition. The experiments are conducted on a flat plate at velocities of 8-14 m/s and the transition of the laminar boundary-layer to turbulence is forced by a disturbance source which is turned on and off with frequencies up to 10 Hz. Three different heating techniques are used to apply the required difference between fluid and structure temperature: a heated aluminum structure is used as an internal structure heating technique, a conductive paint acts as a surface bounded heater, while an IR heater serves as an example for an external heating technique. For comparison of all heating techniques a normalization is introduced and the frequency response of the measured IR camera signal is analyzed. Finally, the different heating techniques are compared and consequences for the design of experiments on laminar-turbulent transition are discussed.

1. Introduction Infrared (IR) thermography is a common measurement technique in many research fields (Astarita & Carlomagno, 2012; Carlomagno & Cardone, 2010; Meola & Carlomagno, 2004) and nowadays a widely used tool in wind-tunnel environments (Gartenberg & OBERTS, 1992; Le Sant, Marchand, Millan, & Fontaine, 2002). The emitted IR radiation of a wind-tunnel model can be measured for the investigation of several flow phenomena, such as separation or laminar-turbulent transition of boundary layers. The increased heat transfer in turbulent compared to laminar boundary layers results in a surface temperature change during laminar-turbulent transition, provided that the fluid and the surface temperatures are different. The experiments reported in literature have mostly been steady measurements (De Luca, Carlomagno, & Buresti, 1990; Gartenberg & OBERTS, 1991; Quast, 1987) or quasi-steady experiments in rotating systems (Lang, Gardner, Mariappan, Klein, & Raffel, 2015; Richter & Schülein, 2014). Even for in-flight experiments IR thermography is an excellent non-intrusive measurement technique and allows a rather uncomplicated visualization of flow phenomena (Banks, vanDam, Shiu, & Miller, 2000;


Crawford, Duncan Jr, West, & Saric, 2014; Horstmann, Quast, & Redeker, 1990). Recent approaches by (Crawford, Duncan Jr, West, & Saric, 2015; Raffel & Merz, 2014; Richter & Schülein, 2014) show that there is large potential for an optimized image post-processing, especially regarding IR thermography measurements, which allow the measurement of unsteady laminar-turbulent transition. Image subtraction techniques have been applied i.e. by (Gartenberg & Wright, 1994) and recent measurements on a helicopter fuselage model by Grawunder illustrate the potential of image processing algorithms for dynamic detection of transition. Nonetheless, the design of the aerodynamic surface, its thermal response time and the employed heating techniques are of major importance for successful dynamic experiments involving transition. Before introducing the experimental setup and measurement equipment in section 2 the fundamental relations and boundary-layer mechanisms required for the transition detection by IR thermography are discussed in section 3. The experimental procedure, including information on the base flow and data processing methods are presented in section 4 while the comparison of the different heating techniques is shown in section 5. Section 6 discusses the influence of the cover layer of the aerodynamic surface and concluding remarks summarize the major results in section 7. 2. Transition Detection by Infrared Thermography There are different mechanisms which lead to laminar-turbulent transition, as discussed in a comprehensive review by (Saric, Reed, & Kerschen, 2002). The mechanism which is most important for the measurement results presented in this manuscript is Tollmien-Schlichting (TS) wave dominated transition. TS wave dominated transition is a common transition scenario on an unswept wing with a smooth surface as it occurs, for example, on a glider wing (Fig. 1). This transition mechanism is well understood and can usually be treated as a 2D problem. Cross-flow transition is the dominant transition mechanism for most modern transport aircraft as it is caused by cross-flow vortices developing at the swept wing leading edge (Saric, Reed, & White, 2003). Wind tunnel and flight experiments on cross-flow dominated laminar-turbulent transition with IR thermography have been conducted e.g. by (Crawford, Duncan Jr, West, & Saric, 2013). The term bypass transition summarizes transition scenarios, which cannot be assigned to one of the two scenarios described above. It is caused by insects, surface roughness, external sound or other disturbance sources, in particular arising from free-stream turbulence.


Fig. 1 Laminar-turbulent transition on a wing.

IR thermography for detecting the laminar-turbulent transition is a thermal measurement technique and relates to the forced convection on the considered body surface. The convective heat transfer 𝑞can be calculated with a heat transfer coefficient ℎ, the adiabatic wall termperature 𝑇% (or fluid temperature for low Mach numbers) and the wall temperature 𝑇&:

𝑞'()* = ℎ(𝑇& − 𝑇%) (1) ℎ = /

0𝑐2𝑈%

456

(2)

where ℎ is closely related to the friction coefficient 𝑐2, the free-stream velocity𝑈%, the thermal conductivity of the fluid 𝑘2, and the kinematic viscosity 𝜈. While equation (1) is often called Newtons cooling law, equation (2) arises from the Reynold's analogy (Schlichting & Gersten, 2003), relating momentum transfer to heat (energy) transfer in the boundary layer for Prandtl number 𝑃𝑟 = 1, which is a valid assumption for air. Usually 𝑐2 is one order of magnitude higher for the turbulent case compared to the laminar boundary layer state. If the heat transfer to the fluid 𝑞'()* is held constant e.g. by constant electrical heating, and the wall friction 𝑐2 increases due to laminar-turbulent transition, the wall temperature 𝑇& will decrease. The resulting change in wall temperature ∆𝑇& = 𝑇&,>?@ − 𝑇&,ABCD can then be measured e.g. by an IR camera. Gartenberg (Gartenberg & Wright, 1994) discusses this heat conduction issue for conventional wind tunnel models made of metal or composite-based fiberglass/carbon-fiber models. IR thermography has become a standard measurement technique for most large wind tunnel facilities, even though literature about its application is rare. Nevertheless, there exists considerable experience in designing models with aerodynamic surfaces which are suitable for IR thermography. In many cases models with adiabatic walls, mostly composite structures, are already prepared for IR thermography transition measurements. In closed, but especially in


open test section wind tunnels, an external IR heating source can easily be installed. Surface heating with conductive paint is also an option to prepare existing wind tunnel models and could be an option to replace temperature sensitive paint (TSP), if electrical connections and uniform heating can be realized. For steady flow conditions the most important requirement is to reach the “minimum resolvable temperature difference” (MRTD), which is of the order of 50 mK for nitrogen or Stirling cooled quantum detector IR cameras, but strongly depends on the emissivity 𝜀 of the surface. The MRTD is governed by the “noise equivalent temperature difference” (NETD) and the spatial resolution (Vollmer & Möllmann, 2010), both are standardized values. If the difference in surface temperature between laminar and turbulent regions Δ𝑇& is higher than the MRTD the transition location can be captured. In contrast to the steady measurements, unsteady IR thermography requires a fast response time of the surface temperature with respect to the instantaneous boundary-layer state. The dynamic properties of the surface are closely related to the model heating technique as discussed below in section 5. 2.1. Unsteady Heat Transfer Before unsteady heat transfer problems can be discussed the definition of material property coefficients and dimensionless numbers is useful. A comprehensive discussion can be found i.e in the (VDI Heat Atlas, 2010). The ability of a material to conduct thermal energy relative to its ability to store thermal energy is called thermal diffusivity 𝛼:

𝛼 = 4HI'J

(3)

while 𝜌𝑐L is the volumetric heat capacity and 𝑘D the thermal conductivity of the body. For transient or unsteady heat transfer problems the Fourier number 𝐹𝑜 (dimensionless time), which is the ratio of the diffusive heat transfer to the change of stored thermal energy of the object, is defined as follows:

𝐹𝑜 = OAPQ= 4HA

I'JPQ (4)

where 𝑠 is a characteristic length, typically half the thickness for an infinite flat plate heat transfer problem, and 𝑡 is the time.


The Biot number 𝐵𝑖 expresses the ratio of the heat transfer at the aerodynamic surface ℎ to that within the object.

𝐵𝑖 = V4H

(5)

A compact measure for the dynamic behavior of a material is the thermal time constant 𝜏 which can be measured as follows:

𝑇 = 𝑇X + (𝑇 − 𝑇%)𝑒[\] (6)

where 𝑇X is the surface temperature at 𝑡 = 0 and 𝑇% the surface temperature for 𝑡 → ∞. The time constant 𝜏 can be estimated by

𝜏 = I'JPV

(7)

while ℎ can be calculated using equation (2). Besides the required material properties, the local friction coefficient 𝑐2 has to be determined. For a zero pressure gradient boundary layer 𝑐2 can be calculated as follows:

𝑐2,>?@ = X.bbcdef

(8)

𝑐2,ABCD ≈

X.X0hdefi/k

(9)

Transient heat transfer mechanisms are often used for non-destructive testing (NTD) of metal or composite structures and active IR thermography methods like pulse thermography (Jackson, Sherlock, & Moore, 2007) or lock-in thermography (Breitenstein, Warta, Langenkamp, & Thermography—Basics, 2010) are well understood. The requirements discussed in the following go beyond this state-of-the-art and are related to the application for aerodynamic problems. 2.2. Requirements on the Measurement Technique The following main goals can be addressed for the detection of laminar-turbulent transition:

1. High signal-to-noise ratio (SNR) 2. Good dynamic behavior 3. High spatial resolution (minimal smearing)


A high signal-to-noise ratio is dependent on the camera temperature resolution and the surface temperature difference between laminar and turbulent regions (see Fig. 1). Surface heating should be as low as possible, so as not to influence the boundary-layer flow significantly (Stock, 2002). Laminar-turbulent transition does not cause high temperature differences, as they occur for example in cooling applications (Dhungel, Lu, Phillips, Ekkad, & Heidmann, 2009). The rather low difference between fluid and surface temperature also leads to a rather low absolute difference in heat flux between the laminar and turbulent state of the boundary layer. The heat capacity and conductivity of the surface material, as well as the structure below, determines the response time of the whole system. The thermal response time of typical wind tunnel models and wing surfaces is rather high, which makes dynamic measurements challenging. Researchers from DLR Braunschweig and Rzeszow university investigated the thermal response time of different surface structures and material combinations at free convection conditions as a part of the AIM2 project. A selection of the material combinations has been tested in flight to investigate the performance for dynamic IR-Thermography measurements under forced convection and sun radiation heating. The investigations presented in this manuscript focus on the evaluation of different materials and heating techniques under forced convection and certain boundary conditions in a wind tunnel environment. The spatial resolution is dependent on two factors: 1) camera resolution and 2) wall parallel heat conduction. The camera resolution for modern cameras has reached up to 1024x1024px, which is high enough for most applications with a rather small field of view (FOV). The critical factor can be the wall parallel heat conduction, which leads to a smearing effect because the turbulent structures in the boundary layers are not mapped exactly by the surface temperature. Therefore the heat conduction in wall parallel direction should be low in order to avoid loss of spatial resolution due to a smearing effect. Such a smearing effect could occur at regions with strong temperature gradients as shown for the turbulent wedge in Fig. 1. The experimental investigations are conducted on a flat plate in a wind tunnel setup which is described in the following. 3. Experimental Setup


The experimental setup consists of a 1.6 m long flat plate which is horizontally mounted in a 0.45x0.45 m2 closed test section of an open-circuit Eiffel type wind tunnel at the Institute of

Fig. 2 Flat plate setup, equipped with a disturbance source 𝑑, an surface hot-wire sensor 𝑟 and a traversable

boundary-layer hot-wire sensor 𝑏.

Aerodynamics and Fluid Mechanics at TU Darmstadt. A low turbulence intensity of 𝑇𝑢 = 0.1% is measured at the end of the 1:24 contraction nozzle. The following paragraphs describe the flat plate setup in more detail whereas the IR measurement system is presented in section 3.1 and different heating techniques in section 3.2. The flat plate (Fig. 2) consists of a 1:6 elliptical leading edge followed by a middle part and a trailing edge flap which allows adjustment of the base flow to achieve a negligible pressure gradient. The middle part is equipped with several sensors and inserts which are carefully flush mounted in order to ensure laminar flow for the undisturbed case. A disturbance source 𝑑 is positioned at 𝑥q =200 mm downstream of the leading edge. It consists of tubes connected to loudspeakers and mounted below a 0.2 mm wide spanwise slit; further information on the device can be found in (Simon et al., 2015). The disturbance source artificially introduces broad-band, high-amplitude TS-waves and forces laminar-turbulent transition upstream of the surface hot-wire 𝑟, placed at 𝑥C =335 mm. The flush mount design of the surface hot-wire allows monitoring of the boundary-layer state and ensures a laminar boundary layer behind the sensor for the undisturbed case. A 400 mm long exchangeable insert starts from at 𝑥 =455 mm and it contains the part of the aerodynamic surface which is heated during the experiment. 3.1. Infrared Camera


The IR radiation emitted from the flat plate is measured with a FLIR SC3000 Stirling cooled quantum detector (QWIP) infrared camera with a resolution of 80x320px at a frame rate of 150 Hz. The camera is sensitive in the long wave IR spectrum (8 to 9µm) and delivers a NETD of less than 20 mK at 30 °C for an ideal black body. The Stirling cooled detector technology ensures low thermal noise while the non-uniformity corrected (NUC) 14 bit digital output raw signal is streamed online to the measurement PC. With exception of the NUC the camera signal is not calibrated to any physical value like temperature or radiation. The temperature calibration of an IR camera includes many uncertainties like viewing angle, surface material, reflected IR radiation and many others (Vollmer & Möllmann, 2010). The detection of laminar-turbulent transition does not require the knowledge of exact temperature values, but rather only the information whether the flow above the surface is laminar or turbulent. This information is available if the “minimum resolvable temperature difference” (MRTD) is reached (Gartenberg & Wright, 1994). The raw data is well suited for the detection of laminar-turbulent transition as it directly shows the signal intensity 𝐼 (14 bit = 0 to 16383). From now on the surface temperature is not measured as a physical quantity but the signal intensity 𝐼, which is indirectly linked to it. The camera is mounted outside of the wind tunnel because of its large dimensions. The camera lens has an opening angle 30 ° in streamwise direction and 15 ° in spanwise direction. A viewing angle of about 70 °ensures no significant signal loss (Vollmer & Möllmann, 2010) and the focal point is set to the center of the FOV, which has a length of 270 mm in streamwise and an average width of 70 mm in spanwise direction. The IR radiation can pass through a small rectangular hole in the wind tunnel wall. A traversable boundary-layer hot-wire sensor 𝑏 is placed just downstream of the considered FOV to monitor the boundary-layer state and ensure that laminar-turbulent transition does not occur between the sensors 𝑟 and 𝑏. The monitoring ensures stable laminar flow conditions for the undisturbed case and confirms the minor influence of the hole in the wind tunnel wall. 3.2. Surface Heating The application of IR thermography requires a temperature difference between fluid and the aerodynamic surface as discussed in section 2. Three different heating techniques are presented in the following. 3.2.1. External Heating


Heat conduction in the structure of a wing or a wind-tunnel model can cause non-uniformconvection on the surface as the temperature differs spatially even for a laminar or fully turbulent boundary layer state. In addition a high thermal capacity of the model can lead to a

(a) External surface heating. (b) Internal surface heating.

Fig. 3 Flat plate insert, equipped with different surface heating techniques.

heat sink and impair the dynamic thermal response. Therefore a surface modification can be necessary to improve the model properties. For external heating the heat capacity of the structure should be as low as possible and have a high emissivity 𝜀. This leads to a design as it is shown in Fig. 3(a). The structure of the plate is thermally insulated by a foam and covered with a thin cover foil. For the experiments presented in this paper the insulating layer consists of a 3 mm rigid polystyrene foam and one or two layers of a matt black Oracal 631 0.08 mm PVC cover foil. A 300 mm long Heraeus midwave IR emitter is aligned in the streamwise direction and placed outside of the wind tunnel. The nominal power of 500 W can be lowered with a power output regulator. A large percentage of the emitted IR radiation can pass through a Heraeus HSQ 301 quartz glass of 2 mm thickness, which is mounted at the upper wind tunnel wall (Fig. 2). Several in-flight experiments with IR thermography have been conducted on the suction side of a wing (Banks et al., 2000; Riedel, Horstmann, Ronzheimer, & Sitzmann, 1998; Seitz & Horstmann, 2006). Due to the sun radiation the upper wing surface heats up and provides the required temperature difference. In wind tunnel experiments an IR radiator can be used to provide the required heating power, as used for example by (Raffel & Merz, 2014). 3.2.2. Internal Structure Heating


The term “internal heating” describes model heating techniques which are implemented in the model itself. Many wind tunnel models consist of metal and therefore have a high heat capacity. The stored heat can be transferred to the fluid during the experiment if the model temperature is higher than the fluid temperature 𝑇% as it can occur in blow through (Eiffel) wind tunnels. In closed-circuit Göttinger wind tunnels the model can be pre-heated to ensure a temperature difference at the beginning of the experiment. This cooling process of the model leads to an unsteady heat transfer even for steady flow conditions, but can provide sufficient measurement time for some applications until the model temperature approaches the fluid temperature. If steady state heat transfer is desired, in order to enable longer measurements an internal heater has to be used. Conventional resistance heating foils are relatively cheap, flexible and can be adapted to arbitrary configurations. Such a heating foil is used for the setup shown in Fig. 3(b), where an insulating layer below prevents a heat flux into the model structure. The inhomogeneous spatial heating of the foil requires an aluminum sheet of 1.5 mm thickness to evenly distribute the heat. Finally, the aluminum is covered with one or two layers of a matt black Oracal 631 0.08 mm PVC foil to enhance the emissivity 𝜀 and avoid reflections in the IR wave band. 3.2.3. Surface Bounded Heating Besides the resistance heater principle used for the model heating with a heating foil, a conductive paint can be directly painted on the model surface, which avoids the use of a metal sheet for homogenization. If the paint is applied on an insulating material the heat capacity of the model surface is of less consequence. This “surface bounded heating” is the third option which is experimentally investigated as a heating technique. The conductive paint Coating Suisse CF-DJF-3014 is used for the experiments. The paint is directly applied on a 1 mm PVC foam by screen printing. It should be mentioned that the sample size of 60 mm x 60 mm does not allow a flush mount application of the PCV foam on the flat plate but the rectangular piece of foam is put on the cover foil shown in Fig. 3(a). This leads to a forward-facing step of 1 mm upstream of the considered measurement region and therefore a perturbation of the laminar and turbulent boundary layer profiles at this point. Despite this drawback the authors present the measurement results with the conductive paint, as it is a new approach for surface heating for transition detection with IR thermography. The hot-wire sensor 𝑏 is used to document the laminar flow in the measurement region for the undisturbed case. The sample with conductive


paint is connected to the DC power supply with two strips of self-adhesive copper tape, aligned in the streamwise direction. The copper tape is connected to the paint with conductive silver lacquer in order to ensure a proper electrical connection. Another surface bounded heating technique is the application of printed circuit boards, which usually consists of a thin copper layer on top of laminated paper. A similar approach is the use of a thin metallic foil. This approach is beneficial if a water flow is investigated and the IR camera can observe the heat transfer from the back, as done by (Gurka, Liberzon, & Hetsroni, 2004). 4. Experimental Procedure and Data Processing An almost zero pressure gradient is adjusted with the trailing edge flap of the flat plate in order to have a rather low amplified transition scenario and a laminar boundary layer in the considered measurement region of the IR camera. The wind tunnel speed ranges from 𝑈&s = 8 - 14 m/s and different heating techniques are tested. The reference surface hot-wire sensor signals and the movable boundary layer hot-wire sensor signals are acquired for all test runs to ensure a laminar (undisturbed case) or fully turbulent (disturbed case) boundary-layer state in the FOV recorded by the IR camera. For 𝑈&s = 14 m/s sensor 𝑏 still shows a laminar boundary-layer state. Fig. 4 shows the sensor signal 𝑟(𝑡) as well as the disturbance source signal 𝑑(𝑡) for a disturbance source pulse frequency of 𝑓q = 1 Hz at 𝑈&s = 14 m/s. It shows a minimal time shift between the signals because of the downstream distance between disturbance source and sensor, but the signals confirm the fully laminar/turbulent state for each half-cycle.

Fig. 4 Disturbance source signal 𝑑(𝑡) and hot-wire signal 𝑟(𝑡) for a pulsed disturbance source operation at

𝑓q = 1 Hz at 𝑈&s = 14 m/s.

4.1. Experimental Procedure


The IR camera provides 150 IR images per second and a single measurement points can be monitored online with the FLIR ThermaCAM Researcher 2001 software which is also used for data recording. Fig. 5 shows an IR image of the flat plate surface and the heated aluminum sheet. The flow comes from the left and the boundary-layer hot-wire is visible at (𝑥 =140 px, 𝑧 = 20 px). For monitoring purposes a turbulator is placed at (𝑥 =50 px, 𝑧 =70 px).

Fig. 5 IR image of the flat plate surface, the middle part shows the heated aluminum sheet.

The first measurement with one heating technique at a certain 𝑈&s is the cooling process initiated by a sudden change of a laminar to a turbulent boundary layer state. The following intensity curves are time series of an area indicated in Fig. 5 with a black square at (𝑥 =90 px, 𝑧 = 5 px), spatially averaged over three pixels in the spanwise (𝑧) direction. Fig. 6(a) shows the recorded signal intensity 𝐼 for the cooling process. It starts at a steady laminar state when the disturbance source is switched on at 𝑡 =10 s. The IR camera signal intensity 𝐼 drops because the surface temperature suddenly changes after transition to turbulence. It is obvious that the gradient decreases with time and 𝐼 asymptotically tends to a steady state again. The start value 𝐼@?x and end value 𝐼@y) define the maximum achievable difference in signal amplitude 𝐼 for the tested configuration. After recording the reference case shown in Fig. 6(a) the pulse frequency 𝑓q is increased from 1/32 Hz to 10 Hz in nine steps (Table 1). The pulsed disturbance source leads to periodic changes of the IR signal 𝐼 (Fig. 6(b)) which is used for the data processing. Before recording the data of the pulsed mode a quasi-steady state is ensured by waiting until the mean signal intensity does not change anymore. Once the recording is started the measurement time 𝑡@ is adjusted to ensure a minimum of two cycles in the records (Table 1). All measurements have been recorded for 𝑈&s = [8, 12 and 14] m/s.


𝑓q (Hz) 1/32 1/16 1/8 and 1/4 1/2, 1, 2.5, 5, 10 𝑡@ (s) 90 45 30 25

Table 1 Disturbance source pulse frequency 𝑓q and the corresponding measurement time 𝑡@ for each run.

(a) Cooling curve. (b) Pulsed disturbance source operation at 𝑓q =0.25 Hz.

Fig. 6 IR signal intensity 𝐼 for different disturbance source operation modes; internal surface heating (aluminum

sheet) with 1450 W/m2 at 𝑈&s = 14 m/s

The data recorded during pulsed disturbance source operation is used to determine the frequency response of the current setup. It would also be possible to extract information from a single cooling curve (Fig. 6(a)) but the pulsed mode acquisition leads to less noisy results because of the averaging over multiple cycles. Several heating power levels are investigated. A homogeneous temperature distribution over the entire FOV is not easy to establish and the average heating power (W/m2) is only a rough estimation for the conditions at the pixel which is chosen for the data processing. In addition the radiation power of the IR heater is not homogeneous and is partly absorbed by the quartz glass at the wind tunnel wall. Each heater was tested for two to three power levels, therefore the following values of heating power levels are only indications:

• Aluminum sheet [800,1450] W/m2 • IR heater [600,950,1250] W/m2 • Conductive Paint [600,750,1300] W/m2

The different and unprecise heating power levels initially appeared to be a problem, but the data processing can compensate for this due to the linear laws of heat conduction in the boundary layer. 4.2. Data Processing


The aim of the data processing procedure is to reduce the large number of recorded IR images to a frequency response as it is known from other measurement techniques. Fig. 6(b) shows an example of an IR signal for a disturbance source pulse frequency 𝑓q =

/sz=0.25 Hz. A centered

moving average filter with a window length of 5 samples in time is applied to the IR signal to remove high frequency thermal noise. Because of the limited frame rate of 150 Hz the amplitude of the signal can be slightly underestimated for 𝑓q ≥

/sz=5 Hz.

The excitation of the system by the disturbance source is pulsed, but not with a harmonic shape of the excitation. Usually transient cooling processes are characterized by the thermal time constant 𝜏 defined earlier in equation (7). This procedure is possible for investigations with one material, but the chosen setup described above does not lead to exponential cooling curves as indicated in Fig. 9. Because of this the high and low peak values in the time signal must be detected by an algorithm and the average peak-to-peak value 𝐼LL (Fig. 6(b)) is plotted for each 𝑓q. The resulting frequency response is presented in Fig. 7(a). It shows the curve for different power levels of the sample heated with conductive paint. A higher surface temperature leads to a better signal-to-noise ratio, but most important to higher amplitudes between the laminar and turbulent state 𝐼LL. Nevertheless the slope is similar for all three cases and the signal amplitude drops quite quickly for higher 𝑓q. 4.3. Normalized Frequency Response The surface temperature is not homogeneous within the heated surface area and a comparison among the different heating techniques is not possible for the given setup. The linear law for the forced convection (equation (1)) allows the definition of Θ, which is 𝐼LL normalized with the maximum achievable signal difference 𝐼@?x −𝐼@y) (see Fig. 6(a)):

Θ(𝑓q) =}JJ(2z)

}~�f[}~��. (10)


(a) Frequency response. (b) Normalized frequency response.

Fig. 7 Frequency response with conductive paint heater at 𝑈&s = 12 m/s for different heating power levels.

All curves in Fig. 7(b) collapse into one if the normalization is applied. Different heating techniques can now be compared independent of the heating power. It should be noted here that the measured (uncalibrated) signal intensity 𝐼 is not proportional to the actual surface temperatur. Nevertheless the linear approximation for this small temperature range seems to be valid and the different heating techniques can be compared in the following. Similar to the normalization described in equation (10) the time signals can be normalized for better comparison, as e.g. in Fig. 8(b):

𝐼∗ = } A [}~��}~�f[}~��

. (11)

5. Comparison of Different Heating Techniques The three different heating technique approaches are compared: external (IR heater with one layer of cover foil), internal (aluminum sheet with two layers of cover foil) and surface bounded (conductive paint). The frequency response curves in Fig. 8(a) exhibit large differences in amplitude Θ as well as in shape and gradient for different parts in the spectrum. It is obvious that the external IR heater shows a much better performance than the internal technique (aluminum sheet) for frequencies up to 𝑓q =1 Hz. The aluminum sheet performs better at frequencies above 𝑓q =1 Hz compared to the IR heater. The surface bounded heating principle of the conductive paint lies in-between the other two approaches for low frequencies but dominates for 𝑓q ≥0.5 Hz. A closer look at the normalized cooling curves in Fig. 8(b) reveals the slow cooling behavior of the aluminum sheet compared to the other two curves, which is dominated by the high heat capacity of the 1.5 mm thick metal sheet. It takes about 𝑡 =120 s until the wall temperature has reached a steady state again, which explains the poor performance for lower frequencies. The fraction 𝑓' =

/A�

would mark the frequency, where the frequency response curve becomes flat. IR

heater and the conductive paint perform much better in this region due to the low thermal capacity of the heated structure or the paint itself, which is applied on an insulating PVC foam.


(a) Frequency response curves. (b) Cooling curves, experimental data.

(c) Cooling curves, theoretical curve 𝐼∗ 𝑡 = 𝑒[A/�plotted on experimental data.

Fig. 8 Comparison of different heating techniques at 𝑈&s = 14 m/s and at the following heating power levels:

IR heater (1250 W/m2), conductive paint (600 W/m2) and aluminum sheet (800 W/m2).

For the following comparison to unsteady heat transfer theory the material properties are taken from Table 2. If a characteristic time of 0.1 s (10 Hz) is considered the 𝐹𝑜 for aluminum and PVC results in 𝐹𝑜?>B = 3.7 and 𝐹𝑜�� = 1.25. The heat transfer coefficient ∆ℎ is calculated for the given problem with ∆𝑐2 = 𝑐2,ABCD − 𝑐2,>?@. A free stream velocity of 𝑈% = 14 m/s, the conduction coefficient 𝑘2 = 2.5 ∗ 10[0 W/(m K) and a viscosity of 𝜈 = 1.5 ∗ 10[� m2/s lead to ∆ℎ = 38 W/K. The thermal time constant 𝜏can then be calculated with equation (7):

τ?>B = 96 s and τ�� = 2.5 s (12).

𝛼 (m2/s) 𝜌𝑐L (J/m K) s (m) aluminum 8.4 ∗ 10[� 2.4 ∗ 10b 1.5 ∗ 10[�

PVC 8 ∗ 10[� 1.2 ∗ 10b 80 ∗ 10[b Table 2 Material properties for aluminum and PVC.


The theoretical cooling curves for both values of 𝜏 are shown in Fig. 8(c). The curves for the external heating (IR heater) and 𝜏�� match well, the theoretical model seems to be valid. For the internal heating (aluminum sheet) the calculated 𝜏?>B does not match the slope of the experimental curve at all. In order to match the experimental results the theoretical value of 𝜏 is manually set to 𝜏 = 31 s and an offset of -0.14 is added. This shows, that the cover foil significantly influences the cooling behavior of the aluminum sheet and simple theoretical models are not able to predict the behavior well enough. The high frequency performance (𝑓q >0.5 Hz) is mainly dependent on the gradient �}

�A just after

the boundary layer becomes turbulent. This effect of the gradient on 𝐼LL(𝑓q) is clearly visible in the time signals of all three heating techniques for 𝑓q =0.125 Hz and 𝑓q =1 Hz in Fig. 9. A closer look on the left column of figures, which correspond to 𝑓q =0.125 Hz, shows much higher amplitudes for Fig. 9(a) and Fig. 9(e) compared to the aluminum sheet (Fig. 9(c)) as recognized before in the frequency response. Just after switching on the disturbance source the gradient �}

�A

differs significantly for all three cases. The signal drop for the aluminum sheet is much steeper compared to the IR heater, while the conductive paint lies in-between. Later in the cooling process of one cycle there is a transition to a smaller gradient, which is very distinct for the aluminum sheet but rather smooth for IR heater and conductive paint. The influence of the gradient on the performance for higher 𝑓q is particularly visible for a pulse frequency of 𝑓q =1 Hz in the figures on the right hand side of Fig. 9. For 𝑓q >1 Hz the performance of the aluminum sheet is better because the steep gradient allows to follow fast changes of the boundary-layer state, whereas the IR heater shows a zig-zag shaped signal (Fig. 9(b)). The deviating transient behavior is caused by different thermodynamic processes in the cover layer. The IR heater heats up the cover foil itself because of the absorption of the IR radiation. This means the upper layer is heated constantly, independent of the temperature distribution in the cover foil. The foam below the foil avoids any heat loss to the structure, therefore the lower side of the cover foil is assumed to be adiabatic. The lower gradient �}

�A , compared to the aluminum

sheet, is mainly caused by the constant heating.


(a) IR heater, 0.125 Hz. (b) IR heater, 1 Hz.

(c) Aluminum sheet, 0.125 Hz. (d) Aluminum sheet, 1 Hz.

(e) Conductive paint, 0.125 Hz. (f) Conductive paint, 1 Hz.

Fig. 9 Normalized IR signal intensity 𝐼∗ for 𝑓q =0.125 Hz and 𝑓q =1 Hz and three different heating techniques.

The aluminum sheet heats up the cover layer from below. After switching on the disturbance source the heat transfer coefficient ℎ, and therefore also the Biot number 𝐵𝑖, increases because of the turbulent boundary layer state. This causes a drop of the surface temperature, a steep gradient �}

�A and an increased temperature difference in the structure. Because of the higher

temperature difference the amount of transferred heat also rises. Due to the thermal inertia 𝑒 of the cover foil this heat flux from the aluminum to the cover layer is phase shifted and causes the flattening of the curve after the steep drop at the beginning. The flattening leads to smaller amplitudes for lower frequencies 𝑓q <1 Hz in the frequency response curve in Fig. 8(a).


6. Influence of the Cover Layer Thickness The results discussed above confirm that the heat capacity and heat conduction in the model surface material can significantly alter the characteristics of the frequency response curves. Especially the cover foil has been found to play an important role during the design of the experimental setup; therefore the influence of the cover layer thickness is discussed further in the following. Besides the effects on the dynamic behavior the thickness has a major influence on the spatial resolution (see section 2.2). (Crawford et al., 2013) describe an internal heating technique setup and conclude that a rather thick cover layer of 0.3 mm leads to a “much crisper image” compared to a thin cover layer applied on the heated aluminum structure. The investigation of the cover layer thickness on the dynamic behavior in this manuscript is done by simply applying one or two layers of the black Oracal 631 0.08 mm PVC cover foil on the internally or externally heated surface. For an internal heating setup the additional cover layer causes an additional insulation between the aluminum sheet and the fluid (Fig. 3(b)). In contrast, a second cover layer for an externally heated surface leads to a higher heat capacity of the heated layer on the surface (Fig. 3(a)). Both effects are shown in Fig. 10 for a flow velocity of 𝑈&s = 14 m/s and at one power level for each heating technique. An interesting fact is that the additional layer improves the performance of the internal heating setup but impairs it for the external heating setup. For a wide range of 0.125 Hz ≤ 𝑓q ≤2.5 Hz the internal heating setup with two layers of cover foil in Fig. 10(a) shows better performance. The additional layer is an insulation and prevents the fast heat conduction from the aluminum to the upper cover layer. On the other hand the setup with one layer is slightly better for high frequencies (𝑓q >2.5 Hz) because the second layer does

(a) Aluminum sheet. (b) IR heater.

Fig. 10 Frequency response curves at 𝑈&s = 14 m/s and at the following heating power levels:

IR heater (1800 W/m2) and aluminum sheet (1450 W/m2).


not allow a fast heat flow from the aluminum to the upper cover layer. A thicker cover layer therefore leads to an increased performance for low frequencies but acts as a low pass filter. The aluminum sheet alone would lead to a lower Biot number 𝐵𝑖 = V

4H due to the high thermal

conductivity 𝑘D, and the temperature gradient in the metal itself would be low. The cover layer leads to a damping and a phase shift of the transient heat flux between the aluminum sheet and the fluid. A second cover layer increases this effect and causes that the aluminum sheet temperature stays more or less constant for fluctuating ℎ and allows a steeper drop of the foil temperature (see Fig. 9(c)). This effect does not alter the properties for very low frequencies of 𝑓q <0.125 Hz. A thinner cover layer is beneficial for experimental setups with an external heating technique (IR heater). Fig. 10(b) shows the frequency response of the externally heated foil with one and two layers of cover foil. For the whole investigated frequency range the performance is better with only one layer of foil. The reduced heat capacity allows the surface temperature to react faster to changes of the boundary-layer state, because the Biot number 𝐵𝑖 = V

4H decreases.

7. Conclusions The frequency response curves of IR thermography for transition detection have been measured with a pulsed disturbance source on a flat plate. This shows the possibilities of IR thermography for unsteady measurements of laminar-turbulent boundary-layer transition for the first time. Three different heating techniques have been investigated (internal, external, surface bounded). In summary, the external heating technique is particularly well suited if the transition is expected to vary with frequencies of 𝑓q <1 Hz. The low heat capacity of the externally heated cover foil is beneficial for low 𝑓q, but the high heat capacity of the aluminum sheet allows the internal heating technique to react faster 𝑓q >1 Hz. The beneficial properties of both techniques seem to be combined in the surface heating with conductive paint. The conductive paint seems to be advantageous if the model allows its application, but an evenly distributed temperature is not easy to achieve because of the unevenly applied coating. This unevenly distributed temperature distribution can be compensated with a normalization technique (Simon, Schnabel, & Grundmann, 2016), but internal and external heating techniques allow a more homogeneous heating. In combination with gradient postprocessing methods (Gartenberg & Wright, 1994) the internal heating can be much more effective because of the steep gradient �}

�A.


The results of the presented experiments can lead to a better design of wind-tunnel models or even in-flight experiments to investigate effects which lead to an unsteady transition. The effect of the propeller wake of an aircraft or the influence of atmospheric turbulence on laminar-turbulent transition on a wing (Reeh & Tropea, 2015) are only two examples. Acknowledgements The authors gratefully acknowledge the financial support by the German Research Foundation DFG (project GR 3524/4-1). We also wish to thank Klaus de Groot from DLR Braunschweig for the fruitful discussions. Finally the authors thank the “Institut für Luft- und Raumfahrt” (Technische Universität Berlin) for the free loan of the cooled FLIR SC 3000 infrared camera and the IDD (TU Darmstadt) for screen printing of the samples with conductive paint. References Astarita, T., & Carlomagno, G. M. (2012). Infrared Thermography for Thermo-fluid-dynamics: Springer Science

& Business Media. Banks, D. W., vanDam, C., Shiu, H., & Miller, G. (2000). Visualization of in-flight flow phenomena using

infrared thermography. Breitenstein, O., Warta, W., Langenkamp, M., & Thermography—Basics, L.-i. (2010). Use for Evaluating

Electronic Devices and Materials: Springer, Berlin/Heidelberg. Carlomagno, G. M., & Cardone, G. (2010). Infrared thermography for convective heat transfer

measurements. Experiments in fluids, 49(6), 1187-1218. Crawford, B. K., Duncan Jr, G. T., West, D. E., & Saric, W. S. (2013). Laminar-turbulent boundary layer

transition imaging using IR thermography. Optics and Photonics Journal, 3(03), 233. Crawford, B. K., Duncan Jr, G. T., West, D. E., & Saric, W. S. (2014). Quantitative Boundary-Layer

Transition Measurements Using IR Thermography. AIAA SciTech, 13-17. Crawford, B. K., Duncan Jr, G. T., West, D. E., & Saric, W. S. (2015). Robust, automated processing of IR

thermography for quantitative boundary-layer transition measurements. Experiments in fluids, 56(7), 1-11.

De Luca, L., Carlomagno, G., & Buresti, G. (1990). Boundary layer diagnostics by means of an infrared scanning radiometer. Experiments in fluids, 9(3), 121-128.

Dhungel, A., Lu, Y., Phillips, W., Ekkad, S. V., & Heidmann, J. (2009). Film cooling from a row of holes supplemented with antivortex holes. Journal of Turbomachinery, 131(2), 021007.

Gartenberg, E., & OBERTS, A. (1991). Airfoil transition and separation studies using an infrared imaging system. Journal of aircraft, 28(4), 225-230.

Gartenberg, E., & OBERTS, A. (1992). Twenty-five years of aerodynamic research with infrared imaging. Journal of aircraft, 29(2), 161-171.

Gartenberg, E., & Wright, R. E. (1994). Boundary-layer transition detection with infrared imaging emphasizing cryogenic applications. AIAA journal, 32(9), 1875-1882.

Gurka, R., Liberzon, A., & Hetsroni, G. (2004). Detecting coherent patterns in a flume by using PIV and IR imaging techniques. Experiments in fluids, 37(2), 230-236.

Horstmann, K., Quast, A., & Redeker, G. (1990). Flight and wind-tunnel investigations on boundary-layer transition. Journal of aircraft, 27(2), 146-150.

Jackson, C. N., Sherlock, C. N., & Moore, P. O. (2007). Nondestructive testing handbook: infrared and thermal testing: American Society for Nondestructive Testing.


Lang, W., Gardner, A., Mariappan, S., Klein, C., & Raffel, M. (2015). Boundary-layer transition on a rotor blade measured by temperature-sensitive paint, thermal imaging and image derotation. Experiments in fluids, 56(6), 1-14.

Le Sant, Y., Marchand, M., Millan, P., & Fontaine, J. (2002). An overview of infrared thermography techniques used in large wind tunnels. Aerospace science and technology, 6(5), 355-366.

Meola, C., & Carlomagno, G. M. (2004). Recent advances in the use of infrared thermography. Measurement science and technology, 15(9), R27.

Quast, A. W. (1987). Detection of transition by infrared image technique. Paper presented at the ICIASF'87-12th International Congress on Instrumentation in Aerospace Simulation Facilities.

Raffel, M., & Merz, C. B. (2014). Differential infrared thermography for unsteady boundary-layer transition measurements. AIAA journal, 52(9), 2090-2093.

Reeh, A. D., & Tropea, C. (2015). Behaviour of a natural laminar flow aerofoil in flight through atmospheric turbulence. Journal of Fluid Mechanics, 767, 394-429.

Richter, K., & Schülein, E. (2014). Boundary-layer transition measurements on hovering helicopter rotors by infrared thermography. Experiments in fluids, 55(7), 1-13.

Riedel, H., Horstmann, K.-H., Ronzheimer, A., & Sitzmann, M. (1998). Aerodynamic design of a natural laminar flow nacelle and the design validation by flight testing. Aerospace science and technology, 2(1), 1-12.

Saric, W. S., Reed, H. L., & Kerschen, E. J. (2002). Boundary-layer receptivity to freestream disturbances. Annual review of fluid mechanics, 34(1), 291-319.

Saric, W. S., Reed, H. L., & White, E. B. (2003). Stability and transition of three-dimensional boundary layers. Annual review of fluid mechanics, 35(1), 413-440.

Schlichting, H., & Gersten, K. (2003). Boundary-layer theory: Springer Science & Business Media. Seitz, A., & Horstmann, K.-H. (2006). In-flight investigations of tollmien-schlichting waves. Paper presented at

the IUTAM Symposium on One Hundred Years of Boundary Layer Research. Simon, B., Nemitz, T., Rohlfing, J., Fischer, F., Mayer, D., & Grundmann, S. (2015). Active flow control of

laminar boundary layers for variable flow conditions. International Journal of Heat and Fluid Flow, 56, 344-354.

Simon, B., Schnabel, P., & Grundmann, S. (2016). IR Measurements for Quantification of Laminar Boundary Layer Stabilization with DBD Plasma Actuators New Results in Numerical and Experimental Fluid Mechanics X: Contributions to the 19th STAB/DGLR Symposium Munich, Germany, 2014 (Vol. 132): Springer.

Stock, H. W. (2002). Wind tunnel–flight correlation for laminar wings in adiabatic and heating flow conditions. Aerospace science and technology, 6(4), 245-257.

VDI Heat Atlas. (2010). Springer. Vollmer, M., & Möllmann, K.-P. (2010). Infrared thermal imaging: fundamentals, research and applications:

John Wiley & Sons.

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