# On the Flow Behavior in Rotor-Stator System with  Superposed Flow

Authors: Roger Debuchy; Fadi Abdel Nour; Gérard Bois
Journal: International Journal of Rotating Machinery
 (2008)

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## Abstract

The flow between a rotor and a stator at high Reynolds number and small Ekman number is divided into three domains, two boundary layers adjacent to the discs separated by a central core. In the present work, a simple theoretical approach provides analytical solutions for the radial distribution of the core swirl ratio valid for a rotor-stator system with a superposed radial inflow rate. At first, the flow in the rotor boundary layer is assumed to behave as expressed by Owen and Rogers (1989) in the case of a turbulent flow on a rotating single disc. On the stator side, a necessary compensation flow rate must take place according to the conservation of mass. It is found that this compensation flow rate cannot be estimated with a good accuracy using the hypotheses of a stationary disc in a rotating fluid by Owen and Rogers (1989). Thus, two innovative weighting functions are tested, leading to new analytical laws relating the core swirl ratioK to the coefficient of flow rate Cqr  introduced by Poncet et al. (2005). The adequacy between the theoretical solutions and numerous results of the literature is clearly improved and the discussion allows a better understanding of the flow behavior.


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## Body

## 1. Introduction

Problems involving flows between rotating discs could be representative of
a lot of situations often encountered in turbomachinery, such as cooling system
of modern gas turbines, pumps, water and steam turbines, and so forth. A
detailed review of many recent studies is reported in Owen and Rogers [1].The earlier studies are proposed by von Karman [2] who simplified the
equations of motion to nonlinear differential equations by means of axial
symmetry hypothesis for the infinite disc rotating in a quiescent fluid and
by Bödewadt [3] who focused on a stationary disc in vicinity of a rotating
fluid. Batchelor [4] gathered the two problems in the case of a rotor-stator
system and proposed a model based on the presence of two boundary layers
adjacent to the discs, and separated by a central core rotating as a solid
body. This result has been contrasted by Stewartson [5] who modeled the
presence of one boundary layer (the rotor one). Both solutions have been
experimented for different cavity aspect ratios and Reynolds numbers.A lot of numerical and experimental studies dealing with enclosed rotor-stator
systems can be found in the literature. Daily and Nece [6] individuated four
flow regimes corresponding to laminar or turbulent flow with merged
or “separated” boundary layers. Itoh et al. [7] studied the effect of the
rotational Reynolds number on the mean velocity field for shrouded discs. It is
worth pointing out that, from their studies, the region near the rotating axis remains
laminar even for a high rotational rate. Furthermore, the transition from
laminar to turbulent regions occurs first along the stator side and then
laminar and turbulent flows coexist. Other numerous works dealing with the
flow in enclosed systems can be quoted: Stepanoff [8],
Szeri et al. [9],
Dijkstra and van Heijst [10],
Cheah et al. [11], and more recently
Randriamampianina et al. [12]. Of course, influence of external flow and
peripheral geometry of the cavity is not considered.In case of a rotor-stator cavity, finite and open at the periphery, peripheral
and external conditions are very important. Djaoui et al. [13] and more
recently D'Haudt et al. [14] and Debuchy et al. [15] highlighted the
strong influence of the geometrical dimensionless parameter λ defined by the
ratio of the difference between stator and rotor radii and the axial gap of
the cavity, in the case of an isolated system, that is, without
any superposed inflow. The case of a rotor-stator system with a superposed
centripetal inflow has been investigated by Debuchy et al. [16]. For high
Reynolds number, small gap ratio and small Ekman number, it was found that the
flow structure is strongly modified as soon as a very weak inflow is applied,
enhancing the level of the core swirl ratio near the axis. Poncet et al. [17, 18]
have found an analytical law for the radial distribution of the central core
swirl ratio, in very good agreement with numerous experimental and numerical
results.In the present work, a simplified theoretical approach leads to analytical solutions
for the core swirl ratio in case of a rotor-stator system with superposed inflow.
Starting from results of Owen and Rogers [1], the fluid centrifuged by the
rotor wall is assumed to behave as a turbulent flow on a single disc. On the
other side (rotor side), it is shown that the compensation flow rate cannot be
estimated with a good accuracy using the results of Owen and Rogers [1] in
case of a stationary disc in a rotating fluid because a radial circulation of
fluid also takes place in the central core. Thus, the authors have tested
innovative weighting functions leading to new analytical laws relating the core
swirl ratio K to the coefficient of flow rate
Cqr introduced by Poncet [17]. In the discussion, extensive
comparisons with experimental and numerical results are in a very good
agreement and provide a better comprehension of the flow  behavior.


## 2. Theoretical Approach

The flow takes place between a rotor of radiusR, rotating with an angular velocity Ω, and a stationary disc. The axial distance between the two discs is H
(Figure 1). The flow is assumed to be incompressible, steady, and axisymmetric. r
and z are, respectively, the radial and axial coordinates (r=0
at the axis, z=0 on the rotor).Figure 1
Schema of the rotor-stator cavity.The significant dimensionless parameters of the problem are the axial gap ratio of the rotor-stator systemG, the Reynolds number Re, and the combination of both, giving the Ekman number Ek. They are, respectively, defined by (1)G=HR,R⁢e=ΩR2υ,E⁢k=1R⁢eG2. Following the theoretical work of Debuchy et al. [16], it is assumed that (2)G≪1,R⁢e≫1,E⁢k≪1. These later assumptions correspond to a regime with separated boundary layers where convective forces are dominant.Dimensionless and normalized quantities with superscript * are introduced(3)r=Rr*,z=G⁢Rz*,vθ=ΩRvθ*,vr=U0vr*,vz=GU0vz*,p-pa⁢t⁢m=12ρΩ2R2p*. In rotor-stator cavities with superposed inflow, Debuchy et al. [16] have introduced the following order of magnitude of the radial velocity
U0=ΩR.Substituting (3) in the conservation of mass and momentum equations for laminar flow, the authors obtain (4)1r*∂(r*vr*)∂r*+∂vz*∂z*=0;(5)vr*∂vr*∂r*+vz*∂vr*∂z*-vθ*2r*=-12∂p*∂r*+[Ek∂2vr*∂z*2],vr*∂vθ*∂r*+vz*∂vθ*∂z*-vθ*vr*r*=[Ek∂2vθ*∂z*2],vr*∂vz*∂r*+vz*∂vz*∂z*=-121G2∂p*∂z*+[Ek∂2vz*∂z*2]. In the right-hand side of (5), only the dominant viscous terms have been retained and turbulence has been neglected.Taking conditions (2) into account, relations (5) become (6a)vr*∂vr*∂r*+vz*∂vr*∂z*-vθ*2r*+12∂p*∂r*=0,(6b)vr*∂vθ*∂r*+vz*∂vθ*∂z*-vθ*vr*r*=0,(6c)∂p*∂z*=0. Relations (4), (6a), (6b), and (6c) have been solved by Debuchy et al. [16]. Two distinct solutions have been obtained: the first one corresponds to a fluid in solid body rotation inside the central core and is valid in the case of an isolated system, that is, without any superposed flow. The second solution is valid in a rotor-stator system with superposed inflow
q. The velocity components are functions of
r*
and
z*: (7)vr*=-cr*sinαsin2α(z*-12)-br*sin(2α(z*-12)+β),vθ*=cr*sinα(cos2α(z*-12)-cosα)+br*,vz*=-cαsinα(cos2α(z*-12)-cosα),p*=b2+c2-c2r*2-b2r*2. As a consequence, the core swirl ratio K defined by vθ*/r* at mid-gap (z*=1/2)
is (8)K=A0r*-2+B0,A0 being function of
q, and B0
being a constant.In the present work, the authors assume that the order of magnitude of the radial velocity componentU0
is
q⋅(2πRHH)-1, RH
being the radius of the central hub. This hub which exists in the real machineries allows to avoid the singularity at
r=0.This leads to introduce the dimensionless flow rate coefficient defined by(9)R⁢oH=q⋅(2πΩRHRH)-1. This later parameter is linked to the Rossby number Ro=q⋅(2πΩR2H)-1
introduced by Debuchy et al. [16] by
RoH=Ro⋅η-1
with
η=RH/R. η
is assumed to be larger than 10%.The conservation of mass equation (4) is valid again, whereas relations (5) become (10)RoH2vr*∂vr*∂r*+RoH2vz*∂vr*∂z*-vθ*2r*=-12∂p*∂r*+[RoHEk∂2vr*∂z*2],RoH2vr*∂vθ*∂r*+RoH2vz*∂vθ*∂z*-RoH2vθ*vr*r*=[RoHEk∂2vθ*∂z*2],RoH2vr*∂vz*∂r*+RoH2vz*∂vz*∂z*=-121G2∂p*∂z*+[RoHEk∂2vz*∂z*2]. Taking conditions (2) into account,
R⁢oH⁢~⁢1
leads to relations (5), otherwise when
R⁢oH≪1, relation (6c) is valid again while the radial equilibrium (6a) is to be replaced by (11)vθ*2r*=12dp*dr*. The relevant conclusion is that the pressure p*
and tangential velocity vθ*
are independent of
z*, but no analytical solution could be obtained from relation (11) only. The radial and axial velocity components are linked together by the conservation of mass equation (4). The following theoretical development is based on continuity considerations in order to provide additional information about the radial evolution of p*
and
vθ*.On the rotor side, a volume flow rateqR
induced by centrifugal effect takes place inside the boundary layer. When a superposed inflow rate q is extracted, the equation of conservation of mass, using the dimensionless volume flow rates denoted by superscripts * and defined by q*=q/υr, qC*=qC/υr,
and
qR*=qR/υr, is (12)qR*-q*+qC*=0, where qC*
is called the dimensionless compensation flow rate. Noting that the sign of the term
qC*
could not be determined a priori.On the rotor side, the authors assume thatqR*
corresponds to the dimensionless flow rate entrained on a single rotating disc when the flow is turbulent. Following the results of Owen and Rogers, qR*
is proportional to
R⁢er4/5, R⁢er
being the local Reynolds number. On the other side, when considering the stator as a stationary disc in a rotating fluid, the dimensionless flow rate in the case of turbulent flow is proportional to
R⁢eω4/5. Here, R⁢eω
is the local Reynolds number based on the velocity of the rotating fluid in the central core. Introducing the core swirl ratio defined by K=vθ*/r*
outside the boundary layers, the authors obtain R⁢eω=KR⁢er
and consequently (13)qR*∝(r*2R⁢e)4/5,qC*∝K4/5(r*2R⁢e)4/5. The combination of relations (12) and (13) leads to (14)K=[A1r*-13/5+B1]4/5 with A1
being proportional to q, and B1
being a constant.This relation is valid under the assumption that the compensation flow rate behaves as in the boundary layer of the stationary disc in a rotating fluid. The first term in the right side of relation (14) is linked to the presence of a superposed flow because A1
is proportional to the flow rate q. Following Poncet et al. [17] and introducing the local coefficient of flow rate
Cqr=qR⁢er1/5/(2πΩr3), relation (14) becomes (15)K=[a1Cqr+B1]5/4 with a1
and B1
being constants.As soon as there is a radial circulation of fluid outside boundary layers, which may occur when there is a superposed radial inflow, the second relation of (13) is no longer valid. Consequently, the authors propose a correction by using a weighting function, so that (13) is to be replaced by (16)qR*∝(r*2R⁢e)4/5,qC*∝WK4/5(r*2R⁢e)4/5. Since the flow behavior in a rotor-stator system is closely linked to the level of the core swirl ratio and to the local coefficient of flow rate, the authors foresee to test two types of weighting function.The first one isW=Kc. Starting from Owen and Rogers results [1] in case of laminar flow, hypotheses (13) have to be written with a power of 1/2 instead of 4/5. The combination of relations (12) and (13) leads to the following analytical relation:
K=[A1r*-2+B1]2. This solution, which is widely different from (8), cannot be in good agreement with the experimental results of Debuchy et al. [16]. When introducing a weighting function of the form
W=Kc, the compensation dimensionless volume flow rate becomes qC*∝Kc+1/2(r*2R⁢e)1/2
and solution (8) can be obtained with c=1/2. According to previous remark, the authors foresee that introducing a weighting function of the form W=Kc
in the second relation of (16) should also improve assumption (13) in case of turbulent flow. Consequently, (14) and (15) are to be replaced by (17)K=[A1r*-13/5+B1]1/c,K=[a1Cqr+B1]1/c with A1
being proportional to q, a1
and B1
being constants.The second weighting function is chosen with the form ofW=fCqr
because of the physical properties of rotor-stator flow: circulation of fluid outside boundary layers must occur when superposed flow rate
q
increases, and/or radial position
r
decreases. The weighting function must also verify fCqr→1
when
Cqr→0
in order to be adequate with the original hypothesis (13). Relations (14) and (15) are to be replaced by (18)K=[A1r*-13/5+B1f(A1r*-13/5)]5/4,K=[a1Cqr+B1f(Cqr)]5/4 with
A1
being proportional to q, a1
and B1 being constants.The adequacy between relations (14), (15), (17), and (18), and numerous theoretical, experimental, and numerical results of the literature will be discussed in the following section.


## 3. Results and Discussion

The validity of the theory is now discussed by making comparisons with the
experimental results of Debuchy [19], Debuchy
et al. [16], and also with the theoretical,
numerical, and experimental works of Poncet [17],
Poncet et al. [18, 20].
All results have been obtained in an annular cavity between a rotating disk
equipped with a central hub and a stationary disk with a central outlet
section in order to ensure a superposed flow. The two experimental setups
have already been detailed in Debuchy [19] and
Poncet [17]. The most significant differences are
located at the periphery of the two systems, with the consequence that it may
influence the inlet flow properties. It has been shown by Debuchy
et al. [15] that if the inlet flow comes from
the external fluid at rest, then the peripheral core swirl ratio (i.e., the
preswirl level) is weak, otherwise the preswirl will be higher. Consequently
the authors are more interested in the description of the peripheral geometry
of the two cavities (Figures 2(a)
and 2(b)).Schema of the peripheral inlet conditions.
(a)
Debuchy [19](b)
Poncet [17]The apparatus used by Debuchy [19] was totally open at the periphery. The rotor, 750 mm
in diameter, was rotating at 1500 rev/min and equipped with a central hub, 180 mm
in diameter. The axial gap of the cavity was fixed to 30 mm
which means that the gap ratio was
0.08. A stationary shroud, 8 mm
in thickness, covered the axial upper side of the stator, which made the stator radius larger than that of the rotor. Consequently, the λ
parameter introduced by Debuchy et al. [15] was fixed to 0.27. Measurements were performed in air at Reynolds number fixed to 1.47×106 (i.e., Ekman number fixed to
1.06×10−4).In the rotor-stator system used by Poncet, an external stationary shroud
near the periphery forced the fluid to recirculate near the rotor with the
consequence that the preswirl level is high. The rotor, 500 mm in diameter,
was rotating up to 600 rev/min. The axial gap of the cavity was adjustable
between 1 and 12 mm. Measurements were performed in water using a laser doppler
anemometry.The main characteristics and dimensionless parameter values retained for the discussion section are listed in Table1. In all cases, the conditions (2) are always verified. The values of λ parameter are to a large extent identical in the two experimental setups. In test case A, RoH is much lower than 1, on the contrary to test cases B and C. However, R⁢oH2 is always lower than 1 (R⁢oH2≈8.5×10−2 is the maximum value for the test case C), which means that relation (11) is valid and that the solutions (14), (15), (17), and (18) are to be tested.Table 1Debuchy [19]Poncet [17]Test case ATest case BTest case CR375 mm250 mmH30 mm9 mmΔR8 mm3 mmη0.240.15λ0.270.33G0.080.036Re1.47×106106Ek1.06×10−47.7×10−4RoH6.95×10−217.4×10−229.2×10−2Cw11 8865 9299 881AirWaterHot wire anemometryLaser doppler anemometryLDAIn the first part of the discussion, the authors evaluate the order of magnitude of the terms neglected in relation (11) by comparison to (6a), that is, the order of magnitude of vr*(∂vr*/∂r*)+vz*(∂vr*/∂z*)
with regard to
vθ*2/r*. The quantity ε is used to define the ratio
(vr*(∂vr*/∂r*)+vz*(∂vr*/∂z*))/(vθ*2/r*).The experimental results of test case A have already been compared to solution (7) with values of b, c, α and β adjusted with the experimental results of Debuchy et al. [16]. The ratio ε is computed from relation (7) with b=0.175, c=0.04,
α=0.3, and β=0.09. The obtained
results are plotted in Figure 3 versus r* and z* for 0.24≤r*≤1.0 and 0.05≤z*≤0.95. Boundary layers have not been considered because the theoretical solution is valid only in the central core: viscous terms in the right-hand of relation (6a), (6b), and (6c) have been eliminated, so that the no-slip conditions
vr*=0 at z*=0 and z*=1, vθ*=r* at z*=0, and vθ*=0 at z*=1 were not satisfied in the theoretical solution. As shown in
Figure 3, the absolute value of the ratio ε is always below 13%, the maximum being located on the stator side for the smallest value of
r*.Figure 3
Axial and radial distribution ofε in test case A.For test cases B and C, which have not yet been compared to solution (7), constants b and c have been adjusted firstly starting from the dimensionless
pressure radial distribution (Figure 4). The authors mention that the dimensionless pressure p* is a little bit shifted in order to satisfy the condition p*=0 at r*=1 instead of the condition p*-p*r*=0.92=0 chosen by Poncet et al. [18]. In a second step, the constants α and β are adjusted from the radial and tangential velocity profiles (Figure 5). The obtained values are mentioned in Table 2. For test cases B and C, the constants adjustment is very difficult because the authors only have few experimental results. However, the tangential velocity levels (Figures 5(a) and 5(b)) and the centripetal flow tendency for the radial velocity profiles at
r*=0.56 (Figures 5(c)
and 5(d)) are in good agreement with
the theoretical solution.Table 2bcαβTest case B0.320.040.250.82Test case C0.450.040.250.93Figure 4
Radial pressure distribution.Tangential (a) and (b)) and radial (c) and (d)) velocity profiles.
(a)
Test case B(b)
Test case C(c)
Test case B(d)
Test case CFigure 6
Core swirl ratio distribution.Using the adjusted constants in Table2, the authors have found a good agreement for the radial distributions of the core swirl ratio, except the peripheral zone where the theoretical preswirl level is slightly underestimated. As for test case A, the radial and axial evolution of the ε ratio is plotted for 0.15≤r*≤1.0 and 0.05≤z*≤0.95 (Figures 7
and 8). It is found again that this ratio is
below 13%.Figure 7
ε axial and radial distribution in the test case B.Figure 8
ε axial and radial distribution  in the test case C.The conclusion of the first part of the discussion is that the simplified radial momentum equation (11) can be used instead of relation (6a) since the ratio ε is small. The Rossby number defined in relation (9) is a significant parameter because ε is small for
R⁢oH2≪1. The analytical solutions (14), (15), (17), and (18) in the theoretical part, based on (9) and (11) can be tested in the following part of this discussion.To begin, the validity of the hypothesis (13) which leads to the solutions (14) or (15) is tested. For this purpose, comparisons between the theoretical solution (15) and
results extracted from Poncet et al. [18] are presented
in Figure 9. Poncet et al. [18]
have found a very good agreement between their analytical law K=2[aCqr+b]5/7-1 and numerous experimental results (with the adjusted
values a≈5.9 and b≈0.63) and also with numerical results based on RSM model (with an adjusted value of a slightly lower: a≈5.3).Figure 9
Core swirl ratio
distribution deduced from (15). Comparison with experimental and numerical results from Poncet [17].Testing relation (15), the best results were obtained for a1≈6.7
and B1≈0.52
in the case of experiments and a1≈6.2
and B1≈0.52
for numerical results. However, the authors found that the agreement is to be improved. The last remark had led the authors to insert a weighting function in (16) and to test relations (17) and (18).Considering the weighting functionW=Kc, the best agreement between relation (17) and the experimental, numerical, and theoretical results of the literature has been found for c=3/2. Consequently, relation (17) is to be replaced by (19)K=[A1r*-13/5+B1]2/3,K=[a1Cqr+B1]2/3.The second relation in (19) is validated using the experimental and numerical results from Poncet et al. [18] after adjustment of constants a1
and B1
(a1≈12.3
and B1≈0.25 for comparison with experiments; a1≈11.2
and B1≈0.25 for
comparison with numerical results). The agreement is much better than that
obtained with relation (15), at least
for Cqr≤0.2 (Figure 10).Figure 10
Core swirl ratio distribution deduced
from (19). Comparison with experimental and numerical results from Poncet [17].The first relation in (19) giving the radial
distribution of the core swirl ratio is validated using the experimental results of
test cases A, B, and C. A very good agreement has been found with A1≈0.1
and B1≈0.03 for
test case A, A1≈0.18
and B1≈0.25
for test case B, and A1≈0.30
and B1≈0.25
for test case C for both the core swirl ratio (Figure 11) and
the dimensionless pressure (Figure 12) distributions.Figure 11
Core swirl ratio distribution.Figure 12
Radial pressure distribution computed with (11).Asc=3/2, the weighting function introduced in (16) is K7/10, which is found to be close to the weighting function K1/2 mentioned in the theoretical approach in case of laminar boundary layer.Large discrepancies are observed in the core swirl ratio distribution forCqr≥0.35 whereas the analytical solution from Poncet [17] is valid for values of Cqr below 0.2 as shown in Figure 13. This may be improved with a weighting function of Cqr in (16), as
previously discussed, or with a more realistic hypothesis concerning the rotor boundary
layer. Indeed, it is well known that the rotor flow rate is centripetal for
core swirl ratio upper than 1: this behavior is not taken into account in
hypothesis (13).Figure 13
Core swirl ratio distribution deduced from (19). Comparison with experimental and numerical results from Poncet [17].When testing the weighting functionW=fCqr,
the best agreement between relation (18) and the theoretical, experimental, and numerical results of the literature has been obtained for
W=eϕCqr. As previously discussed, this weighting function has been chosen because W→1
when
Cqr→0, which means that assumption (13) is also valid in the case of rotor-stator system without any superposed flow. Consequently the relation (18) becomes (20)K=[A1r*-13/5+B1eϕr*-13/5]5/4,K=[a1Cqr+B1eφCqr]5/4ϕ
being proportional to q and
φ
being constants.Comparisons between the theoretical law (20) and the experimental data corresponding to test case A (with A1≈0.085, B1≈0.254, and
ϕ≈0.004), test case B (with A1≈0.143, B1≈0.500, and
ϕ≈0.024), and test case C (with A1≈0.238, B1≈0.500, and
ϕ≈0.04) are
shown in Figures 14 and 15.
The results are in very good agreement.Figure 14
Core swirl ratio distribution.Figure 15
Radial pressure distribution computed with (11).Finally, the validity of the second equation of (20) has been tested over the complete positive range of Cqr. Taking Figure 13 as a reference,
one additional configuration has been considered: it consists of the rotor-stator
system in Figure 2(b) with a central hub
generating another value of post-swirl (i.e., the central core swirl ratio
near the hub). Of sure, the constants in the analytical law need to be
adjusted for each configuration. An excellent agreement between theory and
experiments is shown in Figure 16
for Cqr values up to “0.6.”Figure 16
Core swirl ratio distribution deduced
from (20). Comparison with experimental and numerical results from Poncet [17].


## 4. Conclusion

The aim of this work was to propose a simple theoretical approach in order to obtain a new law relating the core swirl ratioK to the dimensionless flow coefficient Cqr in a rotor-stator system with superposed flow. Starting from the published results of Owen and Rogers, the authors assumed that the fluid on the rotor side behaves as a turbulent flow on a rotating single disc. It is shown that the compensation flow rate on the stator side cannot be described when using the hypothesis of a stationary wall in a rotating fluid by Owen and Rogers but that an innovative weighting function is to be introduced. At first, testing a weighting function with the form of
W=Kc
is interesting because it leads to well-known solutions of the literature. The adequacy between theory and numerous published data is good for a limited range of the coefficient of flow rate. This result is clearly improved with a weighting function with the form of
W=eϕCqr. It leads to a new explicit analytical solution for K valid over a wide range of the Cqr coefficient, knowing that the radial distribution of the core swirl ratio is an interesting quantity for turbomachinery applications.



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