One of the features of this rig is that all of the blade rows have a tip gap. This is done to facilitate the operation of the tangential component of the traverse mechanism. In the experimental setup the tip gap at rest is 0.5mm for the stators and 1mm for the rotor. This was originally faithfully modelled in the CFD mesh. The question of the correct method for modelling the tip gap arose as a result of the apparent mismatch between CFD and the experiment in this region as well as the large discrepancy between measured and predicted torque values.
Changes to the tip gap geometry can be justified in a number of different ways:
The plastic blading has a low flexural modulus and long thin sections can be bent easily
The tip gap was not a strictly controlled dimension given the emphasis on the endwall geometry
The presence of a well in the casing which accommodates the cobra probe head when fully retracted to take measurements near the casing
Measurements of the blading at rest in the test rig confirmed the design tip gaps to within 0.2mm – which is the minimum sinter layer thickness. Some evidence of rubbing was
Rotor Total Total Efficiency
Design Experimental Fine Mesh
Medium Mesh Coarse Mesh Refined Stator Mesh
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S2 R1
X4
found but this is clearly a temporary phenomenon as the blading quickly erodes on contact with the casing. A crude finite element analysis of the rotor blade indicates that the tip gap should actually close slightly towards the trailing edge as a result of the gas forces and centrifugal loading.
This led to the examination of three tip gap geometries: A fully shrouded turbine, the nominal experimental setup and finally one with a tapered tip gap over the rotor, expanding from 1mm at the leading edge to 2mm at the trailing edge as a linear function of axial chord.
Further expansion of the work to investigate increased gaps on the rotor and the introduction of a taper to the tip gaps of the stators resulted in unstable flow solutions.
Attention then turned to the investigation of the effect of the inlet boundary layer profile, and two simulations are provided in Figure 38, one being a turbulent Blasius flat plate profile (White, 2006) for the parallel inlet length and the second the average inlet profile measured at X1 both have constant tip gaps in the rotor geometry.
Figure 38: Comparison of tip gap models at rotor exit
The pitch averaged results of selected parameters are presented against the experimental data in Figure 38 for these models, in this case against the contoured rotor at design speed.
Again the CFD results utilise the k-ω SST turbulence model. In addition a further
Stage Total Total Efficiency 0
Experimental Experimental with plug Shrouded
Constant tip gap Tapered tip gap Experimental Inlet Profile Blasius Inlet Profile
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S2 R1
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measurement was made with the probe well plugged with a ¾ insert, this insert was not used in any of the further experiments presented in this work.
Firstly it is apparent that the plug results in large difference in the experimental data at the tip with smaller differences extending over the span from 30-100%. Below 80% span these differences are limited to less than 2° in relative outlet flow angle and 4m/s in speed. The difference in loss coefficient should not be evaluated at its face value however due to its reliance on average values which are effected by the large changes at the tip, while the efficiency shows no discrepancy below 80% span and this quantity is an indication of the consistency in the pressure profile. Similar differences are visible between annular and contoured experimental results (see Figure 49). In addition Van Den Berg and Bertelrud (1988) suggest that 5-hole probes yield questionable results in highly sheared flows, such as is found in the tip gap region, when compared to non-intrusive laser based techniques, something that is likely to be exacerbated by the need to use the probe in a null yawing mode given the lag in the pressure measurement system. Due to the variation in results near the tip gap, either as a result of the highly sheared flows, exacerbated by the use of the pitch ports only in determining the static pressure; or the presence of the probe well means that all area averaged results presented for the X3 location, immediately behind the rotor will exclude the outer third of span in order to remove this uncertainty from the results.
The presence of the probe well is further acknowledged as having a significant effect on the results presented at all measurement locations, throughout this thesis.
As the tip gap at the rotor trailing edge is increased in the CFD results so the CFD results trend towards the experimental data in the tip gap region, with only small changes evident below 80% span. There remains a large discrepancy in the efficiency predicted versus that measured for the stage, in the order of 10%, some of which (~1%) might be explained by disc windage and bearing losses in the experimental case and the lack of transition modelling in the CFD. Increased stator tip gaps and the pursuit of greater changes to the rotor tip gap was thought to be both impractical (as the solutions proved difficult to converge) and not physical. Hence the use of the tapered rotor tip gap was continued as best practice to cope with the effects of the measurement well. The inlet boundary layer investigations were introduced into the work at a late stage and were not utilised throughout the work. There is clear evidence that the reduced massflow rate in the experimentally derived inlet boundary layer profile reduces the torque and hence the efficiency significantly. There is an average mass flow reduction of 4.2% for all six cases comparing measurements at location X1 to the single point total pressure measurement at location X0, chiefly as a result of the deficit at the casing (Figure 44) where the probe is in the influence of the probe well and measurement points are sparse. However the effect on the pitch averaged profiles (chiefly a reduction in the inflection values at 90% and 50%
span) of the all the quantities shown in Figure 38 is not enough to convince one that this modelling simplification is responsible for the discrepancies between experimental and numerical data profiles.
Few authors in the field have, as yet, shown direct comparisons between CFD and experimentally generated efficiency values, preferring instead to use the relative change between cases and compare these. Abdelfattah and Schobeiri (2010) show a comparison between an unsteady CFD computation across a broad speed range and a experiment on a 3
stage machine where the efficiency differs by between 1 and 4.2%. Gregory-Smith and Crossland (2001) compare computational results from three different sources, both steady and unsteady, where the efficiency varies by as much as 6.4% and there is little consistency in terms of the effects of changing to unsteady solutions. With this backdrop the discrepancy in the absolute value of the efficiency predicted and measured in this study cannot be seen as abnormal and is most likely due to a combination of unknown geometry and boundary conditions, turbulence modelling and unsteady effects, to use Denton (2010)‟s classification as well as uncertainty in the bearing losses in the experimental results as well as the effect of the inlet boundary layer which has not been simulated correctly.