Hussain Nouri
email: hussain.nouri@paris.ensam.fr

Christophe Sarraf
email: christophe.sarraf@paris.ensam.fr.

EXPERIMENTAL INVESTIGATIONS ON THE DESIGN OF A DUCTED COUNTER-ROTATING AXIAL FLOW FANS SYSTEM

An experimental study on counter-rotating axial-flow fans in a ducted-flow configuration was carried out. The fans were designed using an inverse method. The rotation rate of each fan is independently controlled. The distance between the fans can vary from 10 to 50 mm by steps of 10 mm. The efficiency is strongly increased compared to a conventional rotor or to a rotor-stator stage. The study of the variation of the rotation rates ratio show that the system is highly efficient on a wide range of flow-rates and pressure rises. However, the change of the axial distance between rotors does not seem to change the overall performances. This system has thus a very flexible use, with a large patch of high efficient operating points in the parameter space. Further local studies including velocity measurements and wall-pressure fluctuations are needed to better understand the interactions between the rotors and to optimize the system.

NOMENCLATURE

INTRODUCTION

Early studied in the 1930's [START_REF] Lesley | Experiments with a counterpropeller[END_REF][START_REF] Lesley | Tandem air propellers[END_REF], the counter-rotating machines arouse a greater interest in the turbomachinery field, particularly for their potential improvement of the efficiency with respect to conventional machines by recovering more kinetic energy from the front rotor exit-flow [START_REF] Sharma | A study of some factors affecting the performance of a contra-rotating axial compressor stage[END_REF][START_REF] Sharma | A study of aeroacoustic performance of a contra-rotating axial flow compressor stage[END_REF][START_REF] Min | Study on the contra-rotating propeller system design and full-scale performance prediction method[END_REF][START_REF] Shigemitsu | Internal flow measurement with ldv at design point of contrarotating axial flow pump[END_REF]. The first counterrotating machines have appeared in aeronautic and marine applications, in open configuration. Nowadays, these machines with two coaxial axial-flow fans that rotate in opposite direction see a revival of interest in several distinct configurations -open and ducted flows, shrouded or not shrouded rotors-in various subsonic regime applications.

All previous studies concluded that the presence of the rear rotor improves the global efficiency [START_REF] Sharma | A study of some factors affecting the performance of a contra-rotating axial compressor stage[END_REF][START_REF] Sharma | A study of aeroacoustic performance of a contra-rotating axial flow compressor stage[END_REF] and also increases the operating flow-rate range. The counter-rotating systems (CRS) moreover allow to reduce the fans radial compactness [START_REF] Shigemitsu | Internal flow measurement with ldv at design point of contrarotating axial flow pump[END_REF]. A CRS requires more axial spacing compared to one simple fan, but not much more than a rotor-stator stage. Another interesting feature of CRS is that it makes it possible to design axial-flow fans with very low specific speed Ω = ω √ Q

(∆pt/ρ) 3/4 with ω the mean angular velocity, Q the flow rate, ∆p t the total pressure rise and ρ the fluid density.

With such advantages (radial compactness and efficieny improvement), the CRS becomes a very interesting solution and the interactions between the rotors needs to be better understood in order to design highly efficient CRS.

The general aim of the present study on ducted counter-rotating axial-flow fans in subsonic regime is to find out a design method of a CRS by investigating the global and the local performances and the interactions between the two rotors. We first present in § 3 the method that have been used to design the Front and the Rear Rotors. The experimental set-up is presented in § 2. Then we report on the overall performances of the system in § 4. The Counter-Rotating system in its default configuration is studied in § 4.1 and the effects of varying the rotation ratio and the axial spacing between the rotors are presented in § 4.2. We then give some conclusive remarks in § 5.

EXPERIMENTAL SET-UP

The Counter-Rotating System is studied in a ductedflow test rig -AERO 2 FANS-that has been built according to the ISO-5801 standards [START_REF]ISO 5801 -Industrial fans Performance testing using standardized airways[END_REF][START_REF] Sarraf | Experimental study of blade thickness effects on the global and local performances of a controlled vortex designed axial-flow fan[END_REF]. The Fig. 1 shows this test rig. It consists of a cylindrical pipe of inner diameter D = 380 mm. A bell mouth is flushmounted at the inlet of the duct to reduce the energy loss due to fluid friction and flow separation of the inlet flow. The upstream face of FR is at a distance 5D from the pipe inlet. A honeycomb is placed upstream of FR to homogenize the incoming flow. Two brushless PANASONIC A4 motors drive each rotor separatly and are hidden in a casing of diameter 0.33D and of length 0.45D, with a warhead-shape end. For the front motor the honeycomb ensures the binding to the tube. The rear motor is bound to the tube by two rod rows (3 and 5 rods, the first row being at 0.1D from the Rear Rotor). An anti-gyration device made of eight metal sheets of thickness 1.5 mm and length 2D is placed 2D downstream of the CRS. It prevents the outgoing flow from having any rotating component and the static pressure evaluated downstream is more reliable. The static pressure of the axial fan is measured 1D downstream of the anti-gyration device, with an average over four flush-mounted pressure taps. To make the installation more compact, a tube bend of 180 o is placed 1D downstream of the pressure taps. The flow rate is measured with a normalized diaphragm, located 10D downstream of the tube bend and 5D upstream of the pipe outlet. The diaphragm has a diameter of 0.73D. An iris damper -originally used for air flow regulation in ducts-is placed at the exit of the pipe to vary the test-bench hydraulic impedance and thereby to vary the operating point of the studied axial-flow fan. Finally, an axial blower can also be used at the exit of the pipe to widen the explored flow-rate. The static pressure losses from the honeycomb, the motors casings and the anti-gyration device have been measured using this axial blower and have been added to the static pressure rise measurements.

The study focuses on the influence of the gap between the rotors (axial spacing S) as well as on the influence of the speed ratio θ = N RR N F R . The axial spacing vary from S = 10 mm to S = 50 mm by steps of 10 mm. Clear Plexiglas blocks of different thickness are used to change the axial spacing. The reason of using transparent material is to allow forthcoming optical measurements of the velocity field with Laser Doppler Anemometry (LDA) or Particle Image Velocimetry (PIV). Unless specified the default axial spacing is S = 10 mm. Regarding the speed ratio, each rotor is driven separatly so all combinations are possible. The default configuration is θ = 0.9 (see Tab. 1).

Awaiting for the installation of the LDA equipment on AERO 2 FANS, another test-rig (Fig. 2 and[9]) was used to perform LDA measurements on the FR, in free-flow configuration. The mean velocity field in the wake of the fans has been measured with Two-Component LDA, using a DANTEC FlowExplorer system. The method to get the three components of the mean velocity field downstream of the fan relies on the axisymmetry of the mean flow, that is a posteriori checked. The definition of the laboratory Cartesian coordinate system {X; Y ; Z} and of the two measuring planes is pictured in Fig. 2. The measurement is first performed in the {Y ; Z} plane at a certain radial distance r from the axis of rotation, and at an (ax- 2). The measured velocities, in polar coordinates, are thus the radial velocity C r (r, z) and the axial velocity C a (r, z). The measuring volume is then precisely traversed to the same radial and axial distances r and z, but in the {X; Y } plane (point 2 in Fig. . 2). Special care has been given to the alignment of the system. The velocities that are mea-sured at this second point, in polar coordinates, are thus the opposite of the tangential velocity C u (r, z) and the axial velocity C a (r, z) that is measured two times. We then check that the two measurements give the same mean axial velocities: the difference in the two planes is less than 1% everywhere.

DESIGN OF THE ROTORS

The design of the rotors is based on the use of the software MFT (Mixed Flow Turbomachinery) developed by the DynFluid Laboratory [START_REF] Noguera | Design and analysis of axial pumps[END_REF] to which an original method has been added specifically for the design of the Rear Rotor of the Counter-Rotating System.

The design point of the CRS is given in Tab. 1. The system is designed to achieve a total pressure rise ∆p t = 420 Pa at flow-rate Q = 1 m 3 .s -1 for a mean rotation rate around 1900 rpm. That would correspond to a specific speed Ω ≃ 2.46 which is far too low a value for an axial machine. The dimensions of the system, the number of blades for the Front Rotor (FR) and of the Rear Rotor (RR) and their rotation rates are then imposed. The system that is presented here has moreover been designed to have a pure axial exit-flow. In that case the static pressure rise of the CRS -that is the only

= ∆p t -1/2 ρ Q/(πD 2 /4) 2 ≃ 373 Pa.
An iterative procedure is then performed. The pressure rise of the Front Rotor is then arbitrarily chosen and FR is designed and quickly analysed as explained in § 3.1. An estimate of the pressure rise that RR would made is then performed (see § 3.2), based on this analyse. If the total pressure rise of the CRS is not met, the design pressure rise of FR is varied and the calculus are made again. The drawback of this method is that the losses and interactions inbetween the two rotors are not taken into account. The effect of the axial spacing S is also not taken into account and is studied in the present Article.

The geometrical characteristics of the rotor blades obtained with this method are summarized in Tab. 2. Pictures of the Front and Rear rotors are given in Fig. 3.

Design of the Front Rotor

The Front Rotor (FR) was designed as a conventional fan that meets the specifications reported in Tab. 1, i.e. Q = 1 m 3 .s -1 , ∆p tF R = 260 Pa at N F R = 2000 rpm. The hub-to-tip radius ratio is R hub /R tip = 0.293 with tip radius R tip = 187.5 mm. The design is based on the inverse method with simplified radial equilibrium. Rotor is built up from blades of circular-arc camber lines, with NACA-65 profiles clipped at 0.95 x/c. To dimension a new axial machine, the design-point specifications must first be specified (total pressure rise ∆p t , flow rate Q, rotation rate N and radii R tip & R hub ). In a second step, the radial distribution of the circumferencial component of the velocity of fluid, C u2 (r), is imposed at the rotor outlet (the possible types of vortex model: free, constant, forced or mixed). Assuming a perfect fluid and taking into account a rough estimate of the total efficiency of η est = 60% and the equation of simplified radial equilibrium (radial momentum conservation), the velocity triangles at the inlet and outlet of the rotor can then be defined for each radius. The blades can then be defined by the local resolution of an inverse problem considering a two-dimensional flow and searching for the best suited cascade to the proposed velocity triangles. Precisely, this resolution is possible only if we impose the blade number and the local Lieblein's diffusion factor D which has an important influence on the optimization criteria. The behavior of the designed machine resulting from the above method can then be analyzed in order to answer the following questions:

• Is the design-point achieved and what are the values of efficiency locally ?

• What are the characteristics of the machine at the neighborhood of the design point ?

For the imposed speed of rotation, the direct method rules are applied in order to determine the velocity triangles corresponding to each flow discharge. The effects due to real fluid are taken partially into account with the introduction of an axial-velocity distribution which considers the boundary layers at the hub and casing. Thus, we can obtain the characteristics of the machine in the vicinity of the design-point discharge.

In the present case, this direct analysis predicts a mean absolute tangential velocity C u2F R ≃ 9.6 m.s -1 with a radial distribution uniform within ±5% (constant vortex design). The Reynolds number based on the inlet relative velocity varies from 0.6 × 10 5 at the hub to 3 × 10 5 at mid-span and 7 × 10 5 at the tip. LDA measurements were carried out 27 mm downstream of the FR. Please remind that these measurements were performed in a free-flow configuration. The flow, therefore, is not constraint as it could be in a ducted-flow test rig and instead the flow centrifugates in the radial direction (see Fig. 4). The velocity and α profiles are compared to the MFT results (assumed to be at z = 0 mm) in Fig. 5. They show the same trends but do not perfectly match. From the hub to the mid-span, C a values from LDA measurements at z = 27 mm are lower than those from MFT because of the large recirculation close to the hub (see Fig. 4(a)) that MFT does not take into account. From the midspan to the shroud the axial velocity is increased compared to MFT values by virtue of flow-rate conservation. One can notice also that the radial velocity is not zero whereas MFT assumes radial equilibrium. Because of the swirl and the flow centrifugation, the assumption made by MFT is not confirmed at all in freeflow configurations. Concerning the tangential component it is very sensitive to the recirculation in the lower region of the blade: as highlighted in [START_REF] Sarraf | Experimental study of blade thickness effects on the global and local performances of a controlled vortex designed axial-flow fan[END_REF] the tangential and meridional velocity fields are strongly correlated. There is almost zero rotation outside the discharge jet. The 3D structure of the exit-flow thus resembles an annular rotating jet. The maximum of tangential velocity is thus more radially outwards in the LDA than in MFT. Finally, we notice that the magnitude of C u is significant as far as z = 90 mm.

The results from LDA measuremnts, though made in a free-flow configuration, give insight into the interest of a CRS. Firstly, it has been one more time observed that a non negligible energy is contained in the tangential component and could be recovered by the RR. Secondly, previous studies showed that increasing the axial spacing decreases the CRS's noise at the cost of performances [START_REF] Sharma | A study of aeroacoustic performance of a contra-rotating axial flow compressor stage[END_REF], but Fig. 4(b) shows that the kinetic energy that could be recovered from the FR starts to decreases relatively far away which means a good compromise could be found between high performances and relativly low noise. The difficulty would be to have a good estimate of the flow angles along the blade relatively far away from the fan exit to design an adapted RR.

Design of the Rear Rotor

The method used for the design of the Rear Rotor (RR) is to considere the velocity and the flow angle at the trailing edge of the FR blades. Therefore, FR was analysed with MFT to retrieve the axial and tangential velocities (C a2F R = C a1RR and C u2F R = C u1RR respectively) and the angle α 2R1 in the absolute reference frame, at the exit along the blade as shown in Fig. 6.

Using a Matlab script and following an iterative procedure, the RR is drawn in such a way that the exit flow is axial, that is α 2RR = 0 o . The second hypothesis is that the axial velocity profile is kept constant accross RR, i.e. C a2RR (r) = C a1RR (r). Under these asumptions, the total pressure rise of RR should be ∆p tRR = η est ρ U mRR C u2F R ≃ 0.6 × 1.2 × 22.9 × 9.6 ≃ 160 Pa.

In order to completely define the geometry of the blade cascade, we have to compute γ the stagger angle, σ the solidity, c the chord, e % the relative thickness and C z∞0 the camber coefficients. This inverse problem is solved with the following empirical equa-

γ = β 1RR -a, (1)
a is the angle of attack and is obtained by:

a = ∆β RR + 0.94 q(β 1RR ) + 2.07 (2)
with q(β 1RR ) defined by:

q(β 1RR ) = 2.103 -4.01910 -7 β 3.382 1RR (3)
The solidity σ is computed by: Please note that this is a general case where the exit-flow is not constrained to be solely axial

σ -1 = D -1 + C 2RR C 1RR × 2C 1RR |∆C uRR | (4)
with D the Lieblein's diffusion factor. The solidity is computed at the hub and at the tip with D hub = D tip = 0.5 as initial values. Then equation 5 is used to obtain the chord at the hub and the tip:

c = σ 2πR Z ( 5
)
where Z is the number of blades. The chord is computed at the tip and at the hub then the intermediate chords are obtained by linearization. The maximum thickness was set to e max = 6 mm. Finally, the camber coefficients C z∞0 are computed by:

C z∞0 = a + 2.525 p(σ) -0.823 (6)
where p(σ) = 15.535 -12.467e -0.4242σ (7)

We then check that the solidity lays in the range 0.5 ≤ σ ≤ 1.5 and that the camber lays in the range 0 ≤ C z∞0 ≤ 2.7. After several iterations, the RR was drawn with D hub = 0.61 and D tip = 0.46.

RESULTS AND DISCUSSION

Overall performances of the reference sys-

tem ({θ = 0.9 ; S = 10mm})

The characteristics of the Front Rotor rotating alone (RR has been removed from its shaft in that case), of the Rear Rotor rotating alone (FR has been removed) and of the Counter-Rotating system are shown in Fig. 7. The operating speeds are the design speeds, i.e. 2000 rpm for the Front Rotor and 1800 rpm for the Rear Rotor. The static efficiency is defined by equation 8:

η s = ∆P s Q (T F R ω F R ) + (T RR ω RR ) (8)
The nominal flow-rates of the three systems, i.e. the flow-rates at maximum efficiency, are reported in Tab. 3 together with the corresponding static pressure rises and efficiencies.

The Front Rotor rotating alone has a very flat curve (red in Fig. 7). The characteristic curve could not be explored for flow-rates higher than 3800 m 3 .h -1 , i.e. 1.06 m 3 .s -1 , even with the help of the blower. The nominal flow-rate of FR is slightly greater than the design point -it is 3% greater. The measured static pressure rise at the design point is 148 Pa, with a relatively low static efficiency of 45%. This is not surprising with no shroud and a large radial gap of 2.5 mm. Numerical analysis performed with MFT [START_REF] Noguera | Design and analysis of axial pumps[END_REF] and with Fluent 6.3 give very similar results for the static pressure rise (142 ≤ ∆p s ≤ 153 Pa). The total pressure rise predicted by these two different numerical methods -a model using semi-empirical correlations vs Computational Fluid Dynamics-is roughly 260 Pa. The predicted global performances of FR can thus be considered as validated. The remaining question is about the prediction of the exit-flow velocity components, i.e. of the exit-flow angles. This question could of course only be answered with systematic and deep velocity measurements that are scheduled. However, the present global measurements could give partial informations.

The Rear Rotor rotating alone has a steeper curve (blue ⋆ in Fig. 7) and its nominal flow-rate Q ≃ 2600 m 3 .h -1 is lower than the design flow-rate of FR and CRS. This is consistent with the bigger stagger angle of the blades (see Tab. 2).

Let us examine the velocity triangles in Fig. 6 and consider the case with the Front coupled to the Rear Rotor: the incoming velocity C 1RR =C 2F R has an axial component as well as a tangential component. Hence, the flow angle in the relative reference frame Table 3: Nominal points of FR rotating alone at N F R = 2000 rpm, RR rotating alone at N RR = 1800 rpm and CRS at N F R = 2000 rpm and θ = 0.9 (see also Fig. 7

reads: tan(β 1RR ) = U RR + C u1RR C a1RR (9)
Let us consider now the case without the Front Rotor and assume that the flow through the honeycomb is axial. Since the tangential component does not exist anymore and the incoming velocity has only the axial component, equation 9 becomes:

tan(β 1RR ) = U RR C a1RR (10)
Let us now compute the result of equation 9 for the mean radius and at the nominal flow-rate of the CRS, asuming that the tangential velocity is well predicted by MFT, that is U RR ≃ 22.9 m.s -1 , C a1RR ≃ 8.8 m.s -1 and C u1RR = C u2F R ≃ 9.6 m.s -1 . This leads to tan(β 1RR ) ≃ 3.69. If we now suppose that RR rotating alone has its maximum efficiency when the tangent of the inlet flow-angle is equal to this value, equation 10 implies that this is for a flow-rate such that C a1RR = U RR tan( β 1RR ) ≃ 6.2 m.s -1 , i.e. Q ≃ 0.705 m 3 .s -1 or 2540 m 3 .h -1 . This is exactly the nominal flow-rate of RR rotating alone (see Fig. 7 and Tab. 3). The estimations of the angles behind the Front Rotor using the direct analysis of MFT thus seem consistent.

The characteristic curve of the CRS (black • in Fig. 7) is steeper than the characteristic curve of FR. It is roughly parallel to the RR curve. The nominal flow-rate of the CRS matches well with the design flow-rate, i.e. 1 m 3 .s -1 . The static pressure rise at the nominal discharge (∆p sCRS = 340 Pa) is 10% lower than the design point (373 Pa), which is not so bad in view of the rough approximations used to design the system. The CRS has a high static efficiency (η sCRS = 65%) compared to a conventional axial-flow fan or to a rotor-stator stage with similar dimensions, working at such Reynolds numbers [START_REF] Moreau | Efficient stator designed for automotive engine cooling fan systems[END_REF][START_REF] Bakir | Detailed study of an efficient small diameter automotive engine cooling fan system[END_REF]. The gain in efficiency with respect to the Front Rotor is 20 points, whilst an order of magnitude of the maximum gain using a stator is typically 10 points [START_REF] Moreau | Efficient stator designed for automotive engine cooling fan systems[END_REF][START_REF] Bakir | Detailed study of an efficient small diameter automotive engine cooling fan system[END_REF].

The flow-rate range for which the static efficiency lays in the range 60% ≤ η s ≤ 65% is: 2750 The overall performances of the CRS in these conditions are plotted in Fig. 8. As expected, the more the rotation rate of RR increases, the more the static pressure rise of the CRS increases and the nominal flowrate of the CRS increases. The maximal efficiency as a function of θ is plotted in Fig. 9.

Q 4150 m 3 .h -1 ,
For very low rotation rates of RR, i.e. for θ = 0 (Dark green in Fig. 8) and θ = 0.5 (black in Fig. 8), the system is very unefficient: in the first case when the RR is at rest the maximum efficiency hardly reaches 35% which is below the maximal efficiencies of both FR and RR alone. The maximum flow-rate that can be reached is moreover very low in both cases compared to the discharge goal of 3600 m 3 .h -1 .

In the range θ ∈ [0.8 ; 1.2], i.e.

N RR ∈ [1600 ; 2400] rpm, the system is highly efficient. The maximum efficiency increases with θ to reach a maximum value of 66.5% for θ = 1.05 and is then quasiconstant (η s = 66.0% for θ = 1.20). This is a very interesting feature of the Counter-Rotating System. One could imagine, simply by varying the Rear Rotor rotation rate, to work at a constant pressure rise with an efficiency greater than 60% for a large flow-rate range. For instance in the present case, the system could give a constant static pressure rise of 375 Pa with η s ≥ 60% for 3000 ≤ Q ≤ 4250 m 3 .h Figure 10 shows the characteristics curves at the design rotation rates, i.e., N F R = 2000 rpm and θ = 0.9. For this range of axial distances, the overall performances do not vary a lot. There is a slight tendency to a decrease in performance with increasing distances: at the design flow-rate, the difference in static pressure between the best case (S = 10 mm) and the worst case (S = 40 mm) is 17 Pa, which corresponds to a relative decrease of 5%. The efficiency does not vary significantly either. In other studies [START_REF] Sharma | A study of some factors affecting the performance of a contra-rotating axial compressor stage[END_REF][START_REF] Sharma | A study of aeroacoustic performance of a contra-rotating axial flow compressor stage[END_REF] it was reported that the axial spacing had a more significant influence on the overall performances. This is probably due to a difference in the axial distance range.

Let us take as a significant length scale the mean chord length of the Front Rotor (c F R = 58 mm). We introduce the relative axial spacing A:

A = S c F R (11)
The results reported here concern the range A ∈ [0.17 ; 0.86]. The latter value of A is probably too small to see a significant decrease in performances. The effects reported in [START_REF] Sharma | A study of some factors affecting the performance of a contra-rotating axial compressor stage[END_REF][START_REF] Sharma | A study of aeroacoustic performance of a contra-rotating axial flow compressor stage[END_REF] were indeed significant for A = 2. Further studies with greater values of A are scheduled. The effects of the axial spacing on the level of pressure fluctuations will also be investigated in future works.

CONCLUSION

A Counter-Rotating axial-flow fan has been designed according to an iterative method that is relatively fast. It is based on semi-empirical modelization that partly takes into account the losses, boundary layers at hub and casing, and the effects of "low" Reynolds numbers (below 2 × 10 5 ).

The overall performances at the nominal design point are slightly lower than predicted, with a static pressure rise 10% lower. The static efficiency is however remarkably high (η s ≃ 65%) and corresponds to a 20 points gain in efficiency with respect to the Front Rotor maximal efficiency and to a 10 points gain with respect to the Rear Rotor. The overall measurements give first clues that allow to validate the design method.

The Counter-Rotating System has a very flexible use that allows to work at constant flow-rate on a wide range of static pressure rises or to work at constant pressure rise on a wide range of flow-rates, with static efficiency bigger than 60%, simply by varying the Rear Rotor rotation rate. One could thus imagine an efficient closed-loop-controlled axial-flow fan. The overall performances moreover do not significantly vary with the axial spacing in the range A ∈ [0.17 ; 0.86]. This range will be extended to at least 2 Front Rotor chords.

Local measurements of the velocity field in the wake of the Front Rotor rotating alone are scheduled, in order to confirm the design. These measurements will also be of great interest concerning the understanding of the interaction in the space between the rotors. Local measurements of wall-pressure fluctuations are planed, and may show a greater influence of the axial spacing.

The authors finally wish to thank Robert Rey for very fruitfull discussions.

Figure 1 :

 1 Figure 1: Experimental facility for CRS, AERO 2 FANS ial) distance z from the trailing edges of the fan in the Y direction (point 1 in Fig.2). The measured velocities, in polar coordinates, are thus the radial velocity C r (r, z) and the axial velocity C a (r, z). The measuring volume is then precisely traversed to the same radial and axial distances r and z, but in the {X; Y } plane (point 2 in Fig..2). Special care has been given to the alignment of the system. The velocities that are mea-

Figure 2 :

 2 Figure 2: Measurement of the mean exit-flow field. The two measurement planes are displayed with white lines. The origin is centered on the rotation axis in the plane that contains the fans blades trailing edges. The LDA system measures the velocity in the Y and Z directions (point 1: axial C a and radial velocities C r , point 2: axial C z and opposite of the tangential velocities C u

Figure 3 :

 3 Figure 3: Picture of the Front Rotor (left) and Rear Rotor (right)

Figure 4 :

 4 Figure 4: Mean exit-flow measured by LDA at N = 1800RP M with Q = 0.88m 3 .s -1 and brought to N = 2000RP M -Q = 0.98m 3 .s -1 . (a): Mean meridional flow C r (r, z) ; C z (r, z). (b): mean tangential velocity C u . tions that have been validated for NACA-65 cascades [7], for 0.5 ≤ σ ≤ 1.5 and 0 ≤ C z∞0 ≤ 2.7:

Figure 5 :

 5 Figure 5: FR Velocity and angle profiles: (a) axial component, (b) tangential component and (c) absolute reference frame Angle profile. Blue : experimental values (LDA) at z=27 mm at N=1800 RPM brought to 2000 RPM by similitude. Green ⋆: MFT predicted values at z=0 mm. The blade's hub and shroud are delimited by the dashed lines at R=55 mm and R=187.5 mm.

Figure 6 :

 6 Figure 6: Velovity Triangles for the CRS. The fluid is flowing from left to right. Please note that this is a general case where the exit-flow is not constrained to be solely axial

Figure 7 :

 7 Figure 7: Fans characteristics: (a) static pressure rise ∆p s vs flow rate Q; (b) static efficiency η s vs flow rate Q. The axial spacing is S = 10 mm. Red : FR rotating alone at N F R = 2000 rpm (RR has been removed), blue ⋆: RR rotating alone at N RR = 1800 rpm (FR has been removed) and black •: CRS at N F R = 2000 rpm and θ = 0.9. The blue and the dashed lines stand for the design point of the CRS

  that is from 76% of the nominal flowrate up to 115% of the nominal flow-rate. One open question is to what extent the global performances of the CRS are affected by the axial spacing and the speed ratio, and whether the efficient range could be extended by varying the speed ratio.

4. 2

 2 Influence of the rotation ratio θ and of the axial spacing S Rotation ratio θ

Figure 8 :

 8 Figure 8: CRS characteristics at N F R = 2000 rpm, S = 10 mm and θ ∈ [0 ; 1.2] : (a) static pressure rise ∆p s vs flow rate Q; (b) static efficiency η s vs flow rate Q. Dark green : θ = 0, black : θ = 0.5, mustard yellow : θ = 0.8, magenta ▽: θ = 0.85, cyan ⋄: θ = 0.9, red △: θ = 0.95, green * : θ = 1, blue : θ = 1.05, dark green ×: θ = 1.1, black •: θ = 1.15 and purple +: θ = 1.2. The blue and the dashed lines stand for the design point of the CRS In this paragraph, the rotation rate of FR is kept constant at N F R = 2000 rpm, and the rotation rate of RR is varied from 0 rpm

Figure 9 :

 9 Figure 9: Maximal static efficiency η s vs θ for the CRS with N F R = 2000 rpm and S = 10 mm

Figure 10 :

 10 Figure 10: CRS characteristics at various axial spacing: (a) static pressure rise ∆p s vs flow rate Q; (b) static efficiency η s vs flow rate Q. The rotation ratio of FR is N F R = 2000rpm and θ = 0.9. Magenta ▽: S = 10 mm, blue ⋄: S = 20 mm, red △: S = 40 mm and black •: S = 50 mm. The blue and the dashed lines stand for the design point of the CRS Four axial spacings were studied: S = 10, 20, 40 &50 mm.Figure10shows the characteristics curves at the design rotation rates, i.e., N F R = 2000 rpm and θ = 0.9. For this range of axial distances, the overall performances do not vary a lot. There is a slight tendency to a decrease in performance with increasing distances: at the design flow-rate, the difference in static pressure between the best case (S = 10 mm) and the worst case (S = 40 mm) is 17 Pa, which corresponds to a relative decrease of 5%. The efficiency does not vary significantly either. In other studies[START_REF] Sharma | A study of some factors affecting the performance of a contra-rotating axial compressor stage[END_REF][START_REF] Sharma | A study of aeroacoustic performance of a contra-rotating axial flow compressor stage[END_REF] it was reported that the axial spacing had a more significant influence on the overall performances. This is probably due to a difference in the axial distance range.Let us take as a significant length scale the mean chord length of the Front Rotor (c F R = 58 mm). We introduce the relative axial spacing A:

Table 1 :

 1 Design point of the Counter-Rotating System

		CRS	Front Ro-	Rear Ro-
			tor	tor
	D (mm)	380	380	380
	R tip	187.5	187.5	187.5
	(mm)
	R hub	55	55	55
	(mm)
	Z	-	11	7
	∆p t (Pa)	420	260	160
	N (RPM)	1900	2000	1800
	Q	1	1	-
	(m 3 .s -1 )
	Ω	2.46	3.71	-
	Other	Axial exit-	Constant	-
	con-	flow	vortex
	straints
	pressure rise experimentally accessible-should be
	∆p s

Table 2 :

 2 Blade cascade parameters for the two rotors. Radius R (mm). Chord length c (mm). Cascade solidity σ.Stagger angle γ ( o ). Profile designation according to the nomenclature given in Ref.[START_REF] Noguera | Design and analysis of axial pumps[END_REF]: NACA65(xx)yy with (xx) representing the relative camber and yy standing for the relative thickness. Lieblein's diffusion factor D

	Radial position	R	c	σ	γ	profile	D
		Front Rotor (blade thickness 4.5 mm)
	Hub	55	40.3 1.28 23 NACA 65(26)11 0.62
	Mid-span	121.25 58.0 0.84 57 NACA 65(12)07
	Tip	187.5 75.7 0.71 69 NACA 65(07)06 0.44
		Rear Rotor (blade thickness 6 mm)
	Hub	55	58.8 1.18 73 NACA 65(03)10 0.61
	Mid-span	121.25 72.9 0.66 65 NACA 65(05)08
	Tip	187.5 87.1 0.51 75 NACA 65(04)07 0.46

  )

		Front Ro-	Rear Ro-	CRS
		tor	tor
	Maximum	45.28	54.48	64.73
	efficiency
	(%)
	nominal	1.03	0.71	0.99
	Q
	(m 3 .s -1 )
	∆p s (Pa)	148	90	340


Nicolas Lellouche
email: nicolellouche@yahoo.com

MD Frederic Sacher

MD Pierre Jorrot

MD Alain Cariou

MD Christian Spaulding

MD Amandine Aurore

MD Xavier Combes

MD Jerome Fichet

MD Emmanuel Teiger

MD Pierre Jais

MD Jean-Luc Dubois- Randé

MD Michel Haissaguerre

Sudden Cardiac Arrest: ECG Repolarization after Resuscitation

Keywords:

est destinée au dépôt

INTRODUCTION

Out-of-hospital sudden cardiac arrest (OHCA) is a relatively common situation associated with a high mortality rate [START_REF] Myerburg | Sudden cardiac death (structure, function and time-dependence of risk)[END_REF][START_REF] Zipes | Sudden cardiac death[END_REF]. The main cause of OHCA is known to be acute coronary syndrome [START_REF] Atwood | Incidence of EMS treated out-of-hospital cardiac arrest in Europe[END_REF][START_REF] Huikuri | Sudden death due to cardiac arrhythmias[END_REF]. Electrocardiogram (ECG) is the first diagnostic tool that helps in the care of patients with OHCA. However, ECG and particularly repolarization abnormalities are frequent after a cardiac arrest [START_REF] Anyfantakis | Acute coronary angiographic findings in survivors of out-of-hospital cardiac arrest[END_REF] and the prediction of an acute ischemic event from an ECG immediately following cardiac arrest is controversial [START_REF] Anyfantakis | Acute coronary angiographic findings in survivors of out-of-hospital cardiac arrest[END_REF][START_REF] Spaulding | Immediate coronary angiography in survivors of out-of-hospital cardiac arrest[END_REF]. Moreover, early repolarization (ERep) has recently been associated with sudden cardiac death [START_REF] Haïssaguerre | Sudden cardiac arrest associated with early repolarization[END_REF][START_REF] Rosso | J-point elevation in survivors of primary ventricular fibrillation and matched control subjects: incidence and clinical significance[END_REF]. The aim of our study was to assess the value of different repolarization abnormality patterns on the first ECG recorded immediately after OHCA in predicting the presence or absence of acute ischemic disease and the ECG evolution in surviving patients.

METHODS

Patients

The following patients were eligible: age over 18 years old and having experienced OHCA of presumed cardiac origin with at least one ECG in supra-ventricular rhythm.

Exclusion criteria were: cardiac arrest of non-cardiac etiology (i.e. trauma, sepsis, acute respiratory failure. . .); pregnancy, and/or known terminal illness before the occurrence of cardiac arrest.

Study design

This was a retrospective two-center and observational study on OHCA occurring between January 2004 and December 2007. These patients were managed in Henri Mondor University Hospital (Creteil, France) and Cochin University Hospital (Paris, France).

Patient management

In Paris, the management of OHCA involves emergency units based in five hospitals in the city (our two centers are part of these hospitals) and one dispatching center with two physicians. Each emergency unit is equipped with one or more ambulances with resuscitation equipment that are always staffed by one or two physicians trained in emergency medicine. For all patients, cardiopulmonary resuscitation was initiated according to standard procedures. Successfully resuscitated patients were included if the sudden cardiac arrest occurred within one hour of the onset of symptoms in the absence of a clear, non-cardiac cause.

As described in Figure 1, the following diagnostic strategy was applied for every patient: patients who met the criteria were brought directly to the cardiac catheterization laboratory (Cochin or Mondor Hospital). Immediate coronary artery ± left ventricular angiography was performed according to standard techniques. If a recent coronary artery lesion was found, coronary angioplasty and stent implantation were attempted, unless the infarct-related artery was too small or the operator considered the procedure to be technically impossible. Acute coronary lesion was defined as a recent thrombotic coronary occlusion or irregular lesions suggestive of ruptured plaque (Ambrose type 2).

If coronary angiogram was normal, a cerebral CT scan was performed for the diagnosis of acute cerebral hemorrhage. If the cerebral CT scan was normal, a thoracic CT scan was performed to eliminate an acute massive pulmonary embolism which could explain the sudden cardiac arrest. Patients were then transported to the medical intensive care unit.

ECG Analysis

ECG was recorded immediately after out-of-hospital sudden cardiac arrest for all patients using a standard digital recorder with 12 simultaneous leads at a paper speed of 25 mm/s once supra-ventricular rhythm was stable. Two different experts, blinded to the coronary angiogram results and the follow-up, analyzed all ECGs manually. Our study population was divided into 4 groups according to the repolarization abnormalities noted on the first ECG following successful cardiac resuscitation (Figure 2): ERep was defined as an elevation of the QRS-ST junction (J point) in at least two leads. The amplitude of J-point elevation had to be at least 1 mm at 10mm/mV scale (0.1 mV) above the baseline level [START_REF] Klatsky | The early repolarization normal variant electrocardiogram: correlates and consequences[END_REF][START_REF] Mehta | Early repolarization[END_REF], either as QRS slurring (a smooth transition from the QRS segment to the ST segment) or notching (a positive J deflection inscribed on the S wave) in the inferior lead (II, III, and aVF), lateral lead (I, aVL, and V 4 to V 6 ), or both. The anterior precordial leads (V 1 to V 3 ) were excluded from the analysis to avoid the inclusion of patients with right ventricular dysplasia or Brugada syndrome [START_REF] Brugada | Right bundle-branch block and ST-segment elevation in leads V1 through V3: a marker for sudden death in patients without demonstrable structural heart disease[END_REF].

Follow-Up

During the hospital stay, major adverse events, including death from any cause, recurrent cardiac arrest, early reocclusion of the infarct-related artery, early reintervention, and stroke, were recorded. Left ventricular ejection fraction (LVEF) was evaluated by echocardiography within the first days after PCI in all patients using the Simpson method. Patients' neurological recovery and overall performance were estimated at discharge according to current available scales [START_REF]Brain Resuscitation Clinical Trial II Study Group: A randomized clinical study of a calcium-entry blocker (lidoflazine) in the treatment of comatose survivors of cardiac arrest[END_REF]. Briefly, cerebral performance was graded from 1 (good cerebral performance, conscious, alert, able to work and lead a normal life despite possible minor psychological or neurological deficits) to 4 (coma, vegetative state, not conscious, no verbal or psychological interactions with environment) according to Glasgow-Pittsburgh cerebral performance categories [START_REF] Lemiale | Changes in cerebral blood flow and oxygen extraction during post-resuscitation syndrome[END_REF]. Clinical follow-up was obtained in all survivors with a minimum of 3 months after hospital discharge, including evaluation of hemodynamics and clinical data.

Patients with ischemic cardiomyopathy were routinely followed by a cardiologist every 6 months. Patients without structural heart disease or with no acute cause for the cardiac arrest received an implantable cardioverter defibrillator.

For patients who survived, all the ECGs recorded during a follow-up of at least one month after cardiac arrest were analyzed. The presence of ERep during the follow-up was noted according to the previously described criteria. Moreover, for survival patients with unknown initial diagnosis, complementary cardiac exploration was performed, including transthoracic echocardiography, stress test, MRI or an electrophysiological study if necessary.

The study was approved by an institutional review committee and the subjects who survived gave informed consent.

Statistical Analysis

Continuous variables were expressed as mean ± standard deviation (SD), and statistical differences were assessed using the unpaired Student t test or Mann-Whitney U test if necessary. Analysis of variance (ANOVA) was used to compare continuous variables in the 4 groups, each representative of one repolarization abnormality pattern following cardiac arrest. Categorical variables, expressed as numbers or percentages, were compared with the chi-square test or Fisher's exact test. A univariate analysis of variables was performed. All tests were 2-tailed, and a p value <.05 was considered statistically significant.

RESULTS

Population Characteristics

Two hundred and twenty-five patients were included in this study. Overall population characteristics are given in Table 1. Mean age was 59 years old and 79% of our population was male. Hypertension and a preexistent coronary disease were present in 37% and 23% of patients respectively. Fifty-four percent of the cardiac arrest cases were secondary to acute myocardial ischemia. Acute cerebral event was the underlying cause of the cardiac arrest in 3% of cases, pulmonary embolism in 3% and no acute underlying cause was identified in 39% of the patients. Primary recorded rhythm was ventricular fibrillation (VF) in 68%, asystole in 28% and pulseless electrical activity in 4% of cases. Patterns 1, 2, 3, 4 as previously described were observed in 50% (112 patients), 33% (74 patients), 8% (19 patients) and 9% (20 patients) of our population respectively on the first ECG following cardiac resuscitation.

Our population was then divided into 4 groups according to their repolarization pattern.

Patient characteristics according to repolarization pattern

As shown in Table 2, patients with ERep (groups 3 and 4) were significantly younger than other patients. Patients with isolated ST segment elevation (pattern 2) exhibited a higher prevalence of VF as primary rhythm following cardiac arrest compared to the other groups.

Moreover, patients with isolated ST segment elevation were more likely (82%) to have an acute coronary lesion explaining their OHCA, whereas patients with ERep only were more likely to have a non-ischemic cause of death (85%), p<0.0001.

Sensibility and specificity of pattern 2 in predicting acute coronary lesion was 50% and 88% respectively. Positive and negative predictive values were 82% and 60% respectively. Pattern 4 had sensitivity, specificity, positive and negative predictive value of 16%, 98%, 85% and 58% respectively in predicting acutely a non-ischemic event.

Finally, patients with pattern 4 had a significantly higher LVEF (56±10%) following OHCA compared to the other groups (42±16% for pattern 1, 44±15% for pattern 2, 46±11% for pattern 3; p=0.02).

Patient characteristics according to the presence or absence of an acute coronary lesion (Table 3)

In our population, 121 patients had a significant acute coronary lesion to account for OHCA whereas 104 patients did not. Patients with acute coronary lesion responsible for OHCA had a significantly higher rate of ventricular fibrillation (p=0.003) and a lower rate of asystole (p=0.003) compared to patients without acute coronary lesion.

Regarding repolarization abnormalities, these patients exhibited a higher rate of pattern 2 and a significantly lower rate of patterns 1 and 4 compared to other patients without acute coronary lesion. Finally, LVEF following cardiac arrest was similar in both groups.

Characteristics of survival patients and follow-up

In our population of 225 patients, 65 patients (29%) survived after the OHCA episode and were followed as previously described.

As shown in table 4, survival patients were significantly younger (p=0.001), with a lower prevalence of diabetes (p=0.002). Moreover, ventricular fibrillation was more common (p<0.0001) as the primary recorded rhythm after resuscitation and the rate of asystole was lower (p<0.0001) compared to non-survival patients.

As mentioned in Table 4, no etiology was found during initial work-up in 22 survival patients (33%). During follow-up, the precise OHCA diagnosis in this population was: ischemic cardiomyopathy (without acute coronary lesion during OHCA) in 6 patients, right arrhythmogenic dysplasia in 3 patients, hypertrophic cardiomyopathy in 2 patients, coronary spasm in 1 patient, dilated cardiomyopathy in 1 patient, and idiopathic in 9 patients.

No difference was found between survival and non-survival patients regarding the cause of OHCA and the repolarization pattern.

After a mean follow-up of 8±9 months, early repolarization pattern was diagnosed in 7% of patients with pattern 1, 13% with pattern 2, 60% with pattern 3 and 88% with pattern 4 (Figure 3). Concerning the 9 surviving patients without etiology during follow-up, 4 patients had a type 4 pattern at the initial presentation and the same 4 patients displayed ERep during follow-up.

Figure 4 shows examples of ECG after long-term follow-up for survival patients with patterns 3 and 4 immediately after OHCA.

DISCUSSION

Our study demonstrated that immediate repolarization abnormalities following OHCA were associated with specific causes of cardiac arrest: isolated ST segment elevation was associated with acute coronary disease, whereas the presence of ERep was a marker of a non-ischemic event. In the vast majority of patients displaying an ERep pattern at initial presentation, no etiology could be found for cardiac arrest during follow-up. Finally among surviving patients, this pattern if present immediately after OHCA, even though less pronounced, persisted after long term follow-up in most of the patients.

Electrocardiographic ST-segment abnormalities are common after cardiac arrest resuscitation. These repolarization abnormalities can be transient and non-specific [START_REF] Chatterjee | Electrocardiographic changes subsequent to artificial ventricular depolarization[END_REF]. Indeed, these ECG modifications could be caused by electrical cardioversion, electrolyte abnormalities or hypothermia secondary to cardiac arrest [START_REF] Wang | ST-segment elevation in conditions other than acute myocardial infarction[END_REF]. However, specific causes of OHCA such as acute coronary ischemia [START_REF] Anyfantakis | Acute coronary angiographic findings in survivors of out-of-hospital cardiac arrest[END_REF][START_REF] Spaulding | Immediate coronary angiography in survivors of out-of-hospital cardiac arrest[END_REF] or early repolarization could also modify the interpretation of repolarization abnormalities following resuscitation [START_REF] Haïssaguerre | Sudden cardiac arrest associated with early repolarization[END_REF]. Indeed numerous studies have found an excellent correlation between isolated ST segment elevation immediately after OHCA and the presence of acute coronary disease explaining the OHCA [START_REF] Anyfantakis | Acute coronary angiographic findings in survivors of out-of-hospital cardiac arrest[END_REF][START_REF] Spaulding | Immediate coronary angiography in survivors of out-of-hospital cardiac arrest[END_REF]. Consistent with these data, we demonstrated that an isolated ST segment elevation was strongly associated with acute coronary disease, with a positive predictive value of 79%. Thus, we recommend that patients with ST segment elevation immediately after OHCA should go immediately to the cardiac catheterization laboratory for coronary angiogram.

Numerous publications have highlighted the benign nature of ERep [START_REF] Brady | Electrocardiographic manifestations: benign early repolarization[END_REF][START_REF] Klatsky | The early repolarization normal variant electrocardiogram: correlates and consequences[END_REF]. However, recent studies are challenging these findings and indicate that ERep is associated with an increased risk of sudden cardiac death [START_REF] Haïssaguerre | Sudden cardiac arrest associated with early repolarization[END_REF][START_REF] Rosso | J-point elevation in survivors of primary ventricular fibrillation and matched control subjects: incidence and clinical significance[END_REF]. In accordance with this, experimental studies showed that early ERep ECG pattern was secondary to a form of transmural electrical heterogeneity, increasing the risk of malignant arrhythmias [START_REF] Gussak | Early repolarization syndrome: clinical characteristics and possible cellular and ionic mechanisms[END_REF].

Our study showed that isolated ERep was present in 9% of patients after OHCA.

As in Brugada syndrome [START_REF] Probst | Clinical aspects and prognosis of Brugada syndrome in children[END_REF], ERep abnormalities are temporally dynamic and some circumstances such as hypothermia could increase this phenomenon [START_REF] Haïssaguerre | Sudden cardiac arrest associated with early repolarization[END_REF]. One explanation for this high prevalence of ERep could be that as it increases before the onset of malignant arrhythmia, it is also possible that it could increase immediately after the occurrence of cardiac arrest. This hypothesis is confirmed by some observational data collected in our study (Figure 5). Additionally, these patients were mainly diagnosed as having a non-ischemic cardiac arrest (80%).

Eight per cent of our population exhibited ERep associated with ST segment elevation (pattern 3). In this specific subgroup, 47% of patients were found to have a non-ischemic cardiac arrest. The explanation for the association of these two repolarization abnormalities in patients with or without acute coronary lesion is unclear. However, it has been demonstrated experimentally that the pathophysiology of ST segment elevation and the genesis of ERep are similar [START_REF] Shinde | Occurrence of "J waves" in 12-lead ECG as a marker of acute ischemia and their cellular basis[END_REF][START_REF] Yan | Cellular basis for the electrocardiographic J wave[END_REF]. Moreover, consistent with our data, Haissaguerre et al described a dramatic increase in ERep before the onset of malignant arrhythmias, mimicking ST segment elevation during acute myocardial infarction [START_REF] Haïssaguerre | Sudden cardiac arrest associated with early repolarization[END_REF] [START_REF] Haïssaguerre | Ventricular fibrillation with prominent early repolarization associated with a rare variant of KCNJ8/KATP channel[END_REF]. Additionally it has been recently demonstrated that a gene mutation encoding for KATP channels could confer a predisposition to dramatic repolarization changes and ventricular vulnerability in patients with early repolarization [START_REF] Haïssaguerre | Ventricular fibrillation with prominent early repolarization associated with a rare variant of KCNJ8/KATP channel[END_REF]. However, systematic coronary artery angiography should be recommended in this population as no specific parameter could predict the presence of acute coronary lesion.

Finally, during our long-term follow-up 88% of survival patients with isolated ERep immediately after OHCA experienced ERep on further stable ECG, highlighting the specificity of this pattern even under acute circumstances.

Study Limitation

One limitation could be that other factors, such as hypothermia, could have mimicked the presence of ERep immediately after resuscitation. In this case the presence of ERep immediately after cardiac resuscitation could have been overestimated however 88 % have this pattern at follow-up at distance in this group of patients.

Additionally, other repolarization abnormalities such as ST segment depression or isolated T wave changes were not specifically analyzed in this study to try and predict the presence of coronary disease. Finally, autopsy was not performed systematically in the non-survival patients. Other causes of death could therefore have been underdiagnosed, leading to an overestimation in the group with an "unknown" cause of death.

CONCLUSION

In the context of cardiac resuscitation, an ECG with ST elevation favors acute myocardial infarction, whereas the presence of ERep is a marker of a non-ischemic event and future early repolarization syndrome.

Group 1 :

 1 No ST segment elevation nor ERep Group 2: ST segment elevation without ERep Group 3: ST segment elevation and ERep Group 4: ERep only ST segment elevation was defined as an elevation 40 ms after the J point of at least 1 mm in 2 consecutive limb leads and/or 2 mm in two consecutive precordial leads (12).

Figure 1 :

 1 Figure 1: Diagnostic strategy for each patient following successful cardiopulmonary resuscitation. CT= Computed Tomography.

Figure 2 :

 2 Figure 2: Typical ECG examples of pattern 2, 3 and 4 immediately after cardiac resuscitation.

Figure 3 :

 3 Figure 3: Percentage of ERep diagnosed during long-term follow-up according to their initial repolarization pattern in survival patients. ERep: Early repolarization; ECG: electrocardiogram.

Figure 4 :

 4 Figure 4: Example of ECG in survival patients during long-term follow-up.

Figure 5 :

 5 Figure 5: Example of ECG immediately after cardiac resuscitation. Patient with pattern 3 on immediate ECG following cardiac arrest associated with ventricular arrhythmia. The black circle highlights the presence of a notch aspect at the end of the ventricular depolarization. Patient with normal coronary angiogram.