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{{UFR|front=UFR 3-12|description=UFR 3-12 Description|references=UFR 3-12 References|testcase=UFR 3-12 Test Case|evaluation=UFR 3-12 Evaluation|qualityreview=UFR 3-12 Quality Review|bestpractice=UFR 3-12 Best Practice Advice|relatedACs=UFR 3-12 Related ACs}} | {{UFR|front=UFR 3-12|description=UFR 3-12 Description|references=UFR 3-12 References|testcase=UFR 3-12 Test Case|evaluation=UFR 3-12 Evaluation|qualityreview=UFR 3-12 Quality Review|bestpractice=UFR 3-12 Best Practice Advice|relatedACs=UFR 3-12 Related ACs}} | ||
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== Test Case Experiments == | == Test Case Experiments == | ||
The considered test case refers to a high-pressure turbine rotor blade in the VKI short duration isentropic compression tube. In the high-speed, compressible flow environment of this facility wall heat flux measurements were performed with and without film cooling. The comparison with CFD results, however, considers only the data without film cooling. Test facility, model and measurement technique are outlined in Camci and Arts (1985a, 1985b, 1990). For a full description of the cascade geometry reference is to be made to Consigny and Richards (1983). For the experimental set-up blades with a chord length of 80 mm, pitch-to-chord ratio of 0.67 and stagger angle of 38. | The considered test case refers to a high-pressure turbine rotor blade in the VKI short duration isentropic compression tube. In the high-speed, compressible flow environment of this facility wall heat flux measurements were performed with and without film cooling. The comparison with CFD results, however, considers only the data without film cooling. Test facility, model and measurement technique are outlined in Camci and Arts (1985a, 1985b, 1990). For a full description of the cascade geometry reference is to be made to Consigny and Richards (1983). For the experimental set-up blades with a chord length of 80 mm, pitch-to-chord ratio of 0.67 and stagger angle of 38.5° were used. The inlet condition of the flow was the following: Mach number M<sub>in</sub> = 0.25, Reynolds number Re<sub>Cin</sub> = 8.5×10<sup>5</sup>, inflow angle = 30° and the free-stream total temperature T<sub>0∞</sub> = 409.5 K. The Mach number at the exit is M<sub>ex,is</sub> = 0.92. Furthermore, the temperature of the wall was 298 K. The description of the blade geometry can be found in Camci and Arts (1985a, 1985b, 1990). | ||
The ambient turbulence intensity at inlet was equal to Tu<sub>in</sub> = 5% and the length scale was estimated via C<sub> | The ambient turbulence intensity at inlet was equal to Tu<sub>in</sub> = 5% and the length scale was estimated via C<sub>μ</sub><sup>3/4</sup>k<sup>3/2</sup>/ε as ≈ 1 cm. The free-stream turbulence intensity was measured by means of a VKI constant temperature hot-wire anemometer. If Tu<sub>in</sub> would be set to zero, then no noticeable effect from the stagnation point anomaly on heat transfer prediction could be seen. However, turbomachines prevail always a certain amount of ambient turbulence at inlet. | ||
The isentropic Mach number was inferred from static pressure measurements. In Camci and Arts (1985b) the accuracy of the of the pressure reading is given with | The isentropic Mach number was inferred from static pressure measurements. In Camci and Arts (1985b) the accuracy of the of the pressure reading is given with ±0.7% and the one of the temperature with ±0.5%. The local heat flux was deduced from the corresponding time-dependant surface temperature evolution, provided by 45 platinum thin films deposited on the blade surface. The temperature/wall heat flux conversion was obtained from an electrical analogy, see Schultz and Jones (1973), Ligrani et al. (1982). The heat transfer coefficient h<sub>t</sub> is then defined as the ratio between of measured wall heat flux q<sub>w</sub> over the difference between free stream recovery and local wall temperature, T<sub>0∞</sub> − T<sub>w</sub> (recovery factor: 0.896). The flow temperatures were measured using fine tungsten wires and thermocouples. According to Camci and Arts (1985b) the accuracy of the heat transfer coefficient h<sub>t</sub> lies within ±5%. | ||
== CFD Methods == | == CFD Methods == | ||
In order to evaluate the performance of the different turbulence models, Medic and Durbin (2002) used the commercial software package STAR-CD (STAR-CD 1999). This implicit general purpose solver is based on the finite volume concept using a variation of the SIMPLE method (see e.g. Malalasekera & Versteeg, 1996). The turbulence model equations were decoupled form the main flow calculation and solved sequentially. The models are incorporated into the solver via user defined subroutines (v<sup>2</sup>-f and k- | In order to evaluate the performance of the different turbulence models, Medic and Durbin (2002) used the commercial software package STAR-CD (STAR-CD 1999). This implicit general purpose solver is based on the finite volume concept using a variation of the SIMPLE method (see e.g. Malalasekera & Versteeg, 1996). The turbulence model equations were decoupled form the main flow calculation and solved sequentially. The models are incorporated into the solver via user defined subroutines (v<sup>2</sup>-f and k-ω model, respectively). However, the built-in k-ε model was used without any modifications with the exception of the time scale bound which was also incorporated via a subroutine. In the case for the Kato-Launder k-ε model, only the production term P<sub>k</sub> had to be altered. | ||
<font size="-2" color="#888888">© copyright ERCOFTAC 2004</font><br /> | <font size="-2" color="#888888">© copyright ERCOFTAC 2004</font><br /> | ||
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{{UFR|front=UFR 3-12|description=UFR 3-12 Description|references=UFR 3-12 References|testcase=UFR 3-12 Test Case|evaluation=UFR 3-12 Evaluation|qualityreview=UFR 3-12 Quality Review|bestpractice=UFR 3-12 Best Practice Advice|relatedACs=UFR 3-12 Related ACs}} | {{UFR|front=UFR 3-12|description=UFR 3-12 Description|references=UFR 3-12 References|testcase=UFR 3-12 Test Case|evaluation=UFR 3-12 Evaluation|qualityreview=UFR 3-12 Quality Review|bestpractice=UFR 3-12 Best Practice Advice|relatedACs=UFR 3-12 Related ACs}} | ||
Latest revision as of 13:14, 12 February 2017
Stagnation point flow
Underlying Flow Regime 3-12 © copyright ERCOFTAC 2004
Test Case
Brief description of the study test case
Despite several studies about stagnation point flows, a comparison between CFD results and separate experimental data seems to be rare. Some measurements, which unfortunately are not separately published, are found in the paper of Kato and Launder (1993). We have therefore aimed at a test case where sufficient CFD calculations are applied to. This is the case for a VKI experiment, a high-pressure gas turbine rotor blade reported in Camci and Arts (1985a, 1985b, 1990). Medic and Durbin (2002) compare several modifications of the turbulence model with these measurements.
Examined is the transonic compressible flow through a gas turbine blade cascade, where the predicted heat transfer coefficient is of particular interest. The experimental data are compared with the predicted heat transfer coefficient. The turbulent kinetic energy, as a result of different modified two-equation is compared separately.
Test Case Experiments
The considered test case refers to a high-pressure turbine rotor blade in the VKI short duration isentropic compression tube. In the high-speed, compressible flow environment of this facility wall heat flux measurements were performed with and without film cooling. The comparison with CFD results, however, considers only the data without film cooling. Test facility, model and measurement technique are outlined in Camci and Arts (1985a, 1985b, 1990). For a full description of the cascade geometry reference is to be made to Consigny and Richards (1983). For the experimental set-up blades with a chord length of 80 mm, pitch-to-chord ratio of 0.67 and stagger angle of 38.5° were used. The inlet condition of the flow was the following: Mach number Min = 0.25, Reynolds number ReCin = 8.5×105, inflow angle = 30° and the free-stream total temperature T0∞ = 409.5 K. The Mach number at the exit is Mex,is = 0.92. Furthermore, the temperature of the wall was 298 K. The description of the blade geometry can be found in Camci and Arts (1985a, 1985b, 1990).
The ambient turbulence intensity at inlet was equal to Tuin = 5% and the length scale was estimated via Cμ3/4k3/2/ε as ≈ 1 cm. The free-stream turbulence intensity was measured by means of a VKI constant temperature hot-wire anemometer. If Tuin would be set to zero, then no noticeable effect from the stagnation point anomaly on heat transfer prediction could be seen. However, turbomachines prevail always a certain amount of ambient turbulence at inlet.
The isentropic Mach number was inferred from static pressure measurements. In Camci and Arts (1985b) the accuracy of the of the pressure reading is given with ±0.7% and the one of the temperature with ±0.5%. The local heat flux was deduced from the corresponding time-dependant surface temperature evolution, provided by 45 platinum thin films deposited on the blade surface. The temperature/wall heat flux conversion was obtained from an electrical analogy, see Schultz and Jones (1973), Ligrani et al. (1982). The heat transfer coefficient ht is then defined as the ratio between of measured wall heat flux qw over the difference between free stream recovery and local wall temperature, T0∞ − Tw (recovery factor: 0.896). The flow temperatures were measured using fine tungsten wires and thermocouples. According to Camci and Arts (1985b) the accuracy of the heat transfer coefficient ht lies within ±5%.
CFD Methods
In order to evaluate the performance of the different turbulence models, Medic and Durbin (2002) used the commercial software package STAR-CD (STAR-CD 1999). This implicit general purpose solver is based on the finite volume concept using a variation of the SIMPLE method (see e.g. Malalasekera & Versteeg, 1996). The turbulence model equations were decoupled form the main flow calculation and solved sequentially. The models are incorporated into the solver via user defined subroutines (v2-f and k-ω model, respectively). However, the built-in k-ε model was used without any modifications with the exception of the time scale bound which was also incorporated via a subroutine. In the case for the Kato-Launder k-ε model, only the production term Pk had to be altered.
© copyright ERCOFTAC 2004
Contributors: Beat Ribi - MAN Turbomaschinen AG Schweiz