Experimental investigation of gas-liquid separation for two-phase flow within annular duct of an ESP skid

Journal Pre-proof Experimental Investigation of Gas-Liquid Separation for Two-Phase Flow within Annular Duct of an ESP Skid Saon Crispim Vieira, Diogo A.S. Custódio, William Monte Verde, Jorge Luiz Biazussi, Marcelo S. de Castro, Antonio C. Bannwart PII:

S0920-4105(20)31184-0

DOI:

https://doi.org/10.1016/j.petrol.2020.108130

Reference:

PETROL 108130

To appear in:

Journal of Petroleum Science and Engineering

Received Date: 5 June 2020 Revised Date:

18 September 2020

Accepted Date: 11 November 2020

Please cite this article as: Vieira, S.C., Custódio, D.A.S., Verde, W.M., Biazussi, J.L., de Castro, M.S., Bannwart, A.C., Experimental Investigation of Gas-Liquid Separation for Two-Phase Flow within Annular Duct of an ESP Skid, Journal of Petroleum Science and Engineering, https://doi.org/10.1016/ j.petrol.2020.108130. This is a PDF file of an article that has undergone enhancements after acceptance, such as the addition of a cover page and metadata, and formatting for readability, but it is not yet the definitive version of record. This version will undergo additional copyediting, typesetting and review before it is published in its final form, but we are providing this version to give early visibility of the article. Please note that, during the production process, errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain. © 2020 Elsevier B.V. All rights reserved.

Credit author statement

Saon Crispim Vieira: Conceptualization, Methodology, Validation, Investigation, Writing - Original Draft. Diogo A. S. Custódio: Conceptualization, Methodology, Validation, Investigation, Writing - Original Draft. William Monte Verde: Methodology, Writing - Original Draft, Supervision, Project administration.

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Jorge Luiz Biazussi: Software, Methodology, Writing - Review & Editing.

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Marcelo S. de Castro: Conceptualization, Methodology, Writing - Review & Editing. Supervision.

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Antonio C. Bannwart: Writing - Review & Editing, Supervision, Project administration.

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Experimental Investigation of Gas-Liquid Separation for Two-Phase Flow within Annular Duct of an ESP Skid Saon Crispim Vieira¹, Diogo A. S. Custódio², William Monte Verde³, Jorge Luiz Biazussi³, Marcelo S. de Castro², Antonio C. Bannwart² 1

Petrobras – Petróleo Brasileiro S.A School of Mechanical Engineering, University of Campinas, São Paulo, Brazil. 3 Center for Petroleum Studies, University of Campinas, São Paulo, Brazil 2

ABSTRACT

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Although common in several sectors of the industry, undeveloped two-phase flow in annular ducts has yet to receive attention from researchers. This work investigates experimentally the gas-liquid separation phenomenon in the two-phase flows inside annular ducts. The experimental setup enables flow visualization and analysis through high-speed cameras. This experimental setup’s design is based on a non-dimensional and similarity analysis with a real device (Electric Submersible Pump Skid or ESP Skid) in operation at an offshore oilfield in Brazil. The experiments were performed in 5°, 15°, 30° 45° and 60° slopes, horizontally. Flow patterns in the annulus were observed and analyzed. Also, vortices were observed in the intake region. The gas-liquid separation tests enabled the association of the measured separation efficiency with mixture velocities and, consequently, with dimensionless parameters (Froude number) and no-slip void fraction. As observed in the literature, the separation decreases as the mixture velocity increases. The results were qualitatively correlated with literature models, but with high quantitative discrepancies, mainly due to the disregard for radial slip in the models. Also, changes in the trend of separation efficiency were observed and these were caused by flow pattern changes. Keywords: Two-phase Flow, Annular Duct, Intermittent Flow Pattern; ESP Skid; Gas-Liquid Separation. ,

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Nomenclature C0 Distribution Coefficient (-) Bubble diameter (m) Outer diameter (m) Internal diameter (m) Hydraulic diameter (m) Separation efficiency (-) Average separation efficiency (-) Eötvos number (-) Froude number of the mixture (-) Gravity acceleration (m/s2) Axial gravity acceleration component (m/s2) Superficial velocity (m/s) Characteristic length (m) Annular duct 1 length (m) !"# Annular duct 2 length (m) !" Circular duct length (m) Capsule lenght (m) ' Undrained top length (m) ')* Vertical elevation (m) + Morton number (-) , Mass flow rate (kg/s) -. Gauge pressure transducer/measurement (Pa) Volumetric flow rate (m3/kg) Dragged gas flow rate (m3/kg) ,

, ,

V∞TB ∆ $ % & (

Separated gas flow rate (m3/kg) Diameter aspect ratio (-) Reynolds number of the mixture (-) time (s) Temperature transducer (°C) Volume (m3) Volume of dragged gas (m3) Volume of separated gas (m3) Taylor Bubble Velocity (m/s) Void fraction (-) Differential pressure transducer/measurement (Pa) Aspect ratio (-) Inclination angle or slope (°) No-slip gas void fraction (-) Dynamic viscosity (Pa.s) Density (kg/m3) Surface tension (N/m)

Subscripts 1 2

Liquid Phase Gas Phase

m

Mixture

M R

Model Real

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Artificial lifting methods are employed to enable or increase the flow of fluids in oil wells. One of the most important methods is the Electrical Submersible Pump (ESP). In these multiphase flow systems, however, the geometry or equipment operating conditions may lead to fluid segregation and in turn cause instabilities and efficiency degradation of the method. With ESPs, the free gas may lead to performance degradation or even instabilities known as surging (Lea and Bearden, 1982; Monte Verde et al. 2017). The surging phenomenon is characterized by oscillations in flow and head resulting from the operation of centrifugal pumps with gas-liquid two-phase flows. This phenomenon is caused by the dynamic imbalance between the degraded performance curve of the centrifugal pump and the system curve where it operates. When operating in surge condition, the relative velocity of the gas bubbles in relation to the pump impeller is practically zero. This is caused by the flow dynamics in the impeller whose centrifugal field further accelerates the liquid phase, causing the accumulation of the gas phase, which may lead to the total flow interruption, a phenomenon known as gaslocking (Estevam, 2002). Several companies have developed subsea boosting methods called ESP Skid to reduce installation and operational cost. This alternative subsea boosting method is called ESP Skid, presented initially by Rodrigues et al. (2005). More recently, the ESP Skid technology was discussed by Colodette el at. (2007), Teixeira et al. (2012),

Costa et al. (2013), Tarcha et al. (2015, 2016). A Schematic representation of the ESP in the Skid is shown in Figure 1. In ESP Skid, the pumps are housed on the seabed, downstream of the Wet Christmas Tree (WCT) and slightly inclined from the horizontal position. This location results in lower operating pressure, leading to larger instantaneous in situ free gas fractions on the pump intake. Such growth of the free gas fraction tends to intensify gasliquid segregation, increasing the likelihood of surges. Therefore, one must know the dynamics of the fluids separation processes, as such segregation leads to gas accumulation at the top and the configuration of a dynamic level that can fluctuate depending on the operating conditions of the system under analysis (Vieira 2018; Vieira et al. 2020). Lea and Bearden (1982a) pioneered the study of these phenomena by evaluating the effects of natural free gas separation on the performance of ESPs. The authors reported that the separation efficiency decreases with the growth of liquid flow rate and the reduction of the free gas fraction. Alhanati (1993) developed a vertical experimental apparatus to measure the efficiency of natural separation and rotary separators for gas-liquid flows. The experimental results agreed with those of Lea and Bearden (1982a), which showed that the natural separation efficiency decreases with the increase of liquid flow and decrease of liquid-gas ratio.

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1 Introduction

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Figure 1. Schematic representation of the ESP in the Skid.

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Also, Alhanati (1993) proposed a simplified model based on the numerical solution of conservation equations of a two-dimensional drift-flux model and found good agreement with the experimental data, even without considering radial slip. Sambangi (1994) and Lackner (1997) from The Tulsa University Artificial Lift Projects (TUALP), used the same experimental setup developed by Alhanati (1993), expanding the experimental matrix for higher liquid flow rates and mineral oils, respectively. Viloria (1999) developed an experimental apparatus to extend the work of Alhanati (1993) to inclined geometries. An empirical correlation was developed for the void fraction in the intake region, considered different from the void fraction in the annular duct. Marquez and Prado (2001) summarize the experimental conditions used in the cited studies by researchers from TUALP, as shown in Table 1 where the inclination angle is measured from the horizontal. Harun et al. (2001) proposed a mechanistic model to describe natural separation in a vertical well. The model is based on the momentum equation of the mixture, whose domain is a control volume involving the intake region, where a generalist slip law was applied. Experimental data from Alhanati (1993) and Viloria (1999) were used to obtain an empirical correlation for the bubble drag coefficient to close the model.

Liu (2002) and Liu and Prado (2004) proposed a bubble-tracking model to represent the natural separation in a vertical well where, initially, the velocity and pressure field for the liquid phase is solved using a two-dimensional Cartesian coordinate single-phase flow model through stream and vorticity functions. With the velocity and pressure fields, the force balance is solved for a single bubble, whose trajectory is raised over the analysis’ domain. Natural separation is estimated, geometrically, if the bubble trajectory leads to the intake or to the top of the annulus. Additionally, the authors defined the concept of separation circle, which delimits a region of the domain whose bubbles are sucked in for intake. Marquez and Prado (2003) and Marquez (2004) proposed a simplified model for natural separation in a vertical well based on the slip model applied to a control volume in the intake region. The model was closed with classical slip ratio correlations for the vertical direction and a correlation adjusted from experimental data modeled the slip in the radial direction. A simplified model for obtaining the pressure and velocity fields for the liquid phase was also proposed to simplify the bubble tracking approach proposed by Liu (2002). Finally, Marquez (2004) applied stream and vorticity functions approach to the twodimensional mixture model in cylindrical coordinates, numerically solving it to obtain separation efficiencies. Table 2 summarizes the main characteristics of the models

presented in this work for natural separation in annular ducts. The dependence ratios of the natural separation efficiency in relation to the liquid flow rate and the gasliquid fraction suggests that the separation process is based on the balance between the physical drag and buoyancy phenomena. Therefore, the inclination angle, the flow patterns in the annular duct, and force equilibrium in gas bubbles directly impact the phenomenon and, hence, should be discussed and included in a modeling of the separation problem. In none of the cited works was the flow visualized, but the flow patterns were inferred from the classical models

(Barnea, 1987; Taitel et al., 1980) and the separation mechanisms were indirectly elucidated by simulations or by the evaluation of measured parameters such as pressures, flow rates, and efficiencies, so there is a contribution window with regard to the visualization of the phenomena. Finally, the cited works focus on the vertical case (Alhanati, 1993; Liu, 2002; Marquez, 2004), which deals with the inclined case based on a regression of experimental data to calculate the void fraction in front of the intake modified by the radial slip (Viloria, 1999), so there is a lack of experimental data in the inclined case, providing space to contribute to the modeling.

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Table 1. Experimental conditions of studies from TUALP, adapted from Marquez and Prado (2001).

Sambangi (1994)

Lackner (1997)

Viloria (1999)

Air-Water

Air-Water

Air-Mineral Oils

Air-Water

Liquid Flow Rate (bpd)

300; 600; 900

From 1200 to 3000

From 300 to 2700

1000; 1500; 2000

GLR (scf/stb)

50; 100; 200

50; 100; 200; 300

-

Pressure (Psig)

50; 100; 150

re

50; 100; 200; 300

100; 200; 300

100; 200; 300

50; 100; 150

0; 60

0; 30; 60

-

90°

90°

90°, 60° and 30°

-

-

0.5; 0.10; 0.15; 0.20

Fluids

Speed (Hz)

0; 60 90°

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Inclination Angle

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Alhanati (1993)

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Variable

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Experimental conditions

Gas Fraction

-

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Table 2. Characteristics of natural separation models for annular ducts.

Alhanati (1993)

Model Characteristics Vertical water-air flows. Simplified model - numerical solution of a 2D drift-flux model. Good agreement with experimental data. Disregard radial slip.

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Viloria (1999)

Vertical and Inclined water-air flows. Void fraction at the intake correlation proposed, different from the one at the annulus. Disregard radial slip. Inclined flow analysis based on experimental data regression for intake region void fraction.

Harun et al. (2001)

Mechanistic model based on the momentum mixture equation on a control volume in the intake region. Drag correlation from experimental data of Alhanati (1993) and Viloria (1999) Disregard radial slip. No analysis for inclined flows.

Liu and Prado (2004)

Bubble tracking model for vertical flows. Stream and Vorticity functions solved for velocity and pressure of liquid single-phase flow. Calculate force balance on a single bubble with the velocity and pressure fields. Separation given by the bubble trajectory. Separation circle concept. Disregard radial slip. No analysis for inclined flow.

Marquez (2004)

Vertical flow. Simplified model based on slip model. Uses classical slip correlations to close the model. Radial slip modelled based on the experimental data regression. Uses a simplified version of the bubbler tracking model of Liu and Prado (2003). No analysis for inclined flow.

Given the above, this work aims to experimentally study the two-phase flow in the intake of subsea ESP Skid type pumping systems (Costa et al., 2013; Tarcha et al., 2015, 2016). The two-phase flow in the annular duct formed between the capsule and the ESP are evaluated

regarding the fluid segregation phenomenon at the inlet in various inclinations: 5° (similar to the actual equipment), 15°, 30°, 45°, and 60° from horizontal position, using multiphase flow visualization and measurement techniques. For this, an experimental apparatus was

2 Research Methodology The purpose of the experimental apparatus designed in this paper is to reproduce the morphology of the two-phase gas-liquid flow inside the ESP Skid capsule and enable the collection of experimental data for the detailed analysis of the fluid separation process. The experimental campaign was conducted at the Experimental Laboratory of Petroleum (LABPETRO), at the Center for Petroleum Studies (CEPETRO) of the University of Campinas (UNICAMP), Brazil.

= sin5# 6 7 +

(2)

- No-slip void fraction ($): $=

#

+

(3)

where is the volumetric flow rate and the subscripts 1 and 2 represents the liquid and gas phases, respectively. - Separation efficiency ( ): dimensionless number, which is one of the aims of this study, defined by the ratio between the separated gas flow rate ( , ) that migrates to the top of the annular duct, and the inlet gas flow rate ( ).

2.1.1 ESP Skid Capsule Model

=

,

(4)

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The design of the ESP Skid capsule followed a dimensional and similarity analysis from the dimensions and operational conditions of the real equipment. Next, the term "real" represents the equipment installed in the field and the term “model” represents the scaled model, used in the testing setup. The main variables that influence the phenomenon under study were listed, including parameters related to annular duct geometry, PVT and fluid transport properties, and flow characteristics. From Buckingham's π Theorem (Panton, 2013; White, 1998 and 2006), where the variables were inserted into a linear transformation, the main trivial dimensions were found, such as aspect ratio, slope, no-slip void fraction, and separation efficiency, in addition to the main classical dimensionless numbers; which are the Reynolds, Froude, Eötvos and Morton numbers. In this work, the equivalent hydraulic diameter was used, i.e. the difference between the outer and the internal diameter, and mixture properties. The mixture fluid properties are determined from the single-phase gas and liquid properties averaged from the no-slip void fraction. Apart from the multiphase flow, the literature suggests that these dimensionless numbers, written from mixture parameters, are more representative, e.g., the pressure gradient calculation uses friction factors defined as a function of the Reynolds number of the mixture (Wallis, 1969; Shoham, 2005). Therefore, the variables mentioned are defined by: - Aspect ratio ( ): all model linear dimensions are scaled proportional to the real equipment.

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2.1 Experimental Setup

where and are the outer and internal diameters, respectively. is any characteristic linear dimension and the subscribed and , represent the real equipment and the scaled model, respectively. - Slope ( ): trivial dimensionless relative to the duct length and the vertical elevation ( + ).

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designed from a dimensional analysis of the real system. The flow in the annular duct upstream of the intake and the phenomenon of gas-liquid separation at the intake were visualized and experimentally characterized.

- Reynolds Number of the Mixture (

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=

=

,0

,1

=

,0

,1

=

0

1

(1)

&

):

%

(5)

are the density, dynamic viscosity where & , % and and volumetric flow rate of the mixture, respectively, and is the hydraulic diameter, defined by: & = 91 − $<&# + $&

(6)

% = 91 − $<%# + $% and

=

=

#

(7)

+

(8)

−

(9)

- Froude Number of the mixture ( =

):

=

(10)

where is the axial component of the gravity. - Eötvos Number ( ): =

9&# − & < (#

where ( # is the surface tension, and is the gravity. - Morton Number (, ): , =

%#> 9&# − & < &# ( ?#

(11) is the bubble diameter

(12)

Once the fundamental dimensionless numbers that characterize the problem are defined, similarity is used to

<0

(13)

Substituting Eq. (10) into Eq. (13):

@

=A

1

=@

=A

Introducing the aspect ratio ( ) given by Eq. (1) into Eq. (14), the mixture flow rate used in the scaled model testes is: =6 7

,1

=

,0

(15)

<0

7 =6 1

&

%

6 7 = B6

% & 1 7 6 7 6 7C & 1 % 0

(19)

From Eq. (24) it is possible to correlate the properties of the real fluid and the model fluid with the aspect ratio: E

= DB6

& % 7 6 7 C % 1 & 0

(20)

Therefore, once the geometry of the real ESP Skid and its operating condition are known, the aspect ratio is defined and the model dimensions are calculated by Eq. (1) . Using Eqs. (15) and (20) the model operating conditions and the model’s fluid properties are calculated, respectively. Notedly, the no-slip void fractions must be the equal in the model and in real operation. Figure 2 schematically illustrates the ESP Skid geometry. The real equipment is inclined at 5° from the horizontal position in the seabed.

7

0

na

(16)

Substituting Eq. (5) into (16): %

=

(17)

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&

1

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6

<1 = 9

(18)

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Performing the same analyses for the Reynolds number: 9

,0

re

1

% & 1 7 6 7 6 7C & 1 % 0

Substituing Eq. (15) into (18):

(14)

0

= B6

ro

<1 = 9

,1

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9

Again, introducing the aspect ratio ( ) given by Eq. (1) into Eq. (22), the mixture flow rate used in the scaled model testes is:

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obtain the model’s geometry and the operating conditions that provide dynamic similarity to the real equipment installed in the field. The Reynolds and Froude numbers are the most important for proper flow description because they are directly linked to turbulence and fluid segregation. The Eötvos and Morton number will not be similar given the difficulty in controlling parameters such as surface tension and bubble diameter in the model testing. From similarity between the model and real Froude numbers, one obtains:

Figure 2. Simplified Geometry of the ESP Skid capsule. The ESP Skid can be divided into the following sections:

1) Circular Duct: ESP Skid capsule section where there is no internal duct;

Internal Diameter ( )

0.203

0.075

Outer Diameter (

0.296

0,110

0.093

0.035

1.46

1.47

Description Capsule length (

'<

Circular duct length ( Annular duct 1 length ( Annular duct 2 length ( Undrained top length (

Real [m]

Model [m]

21.059

4.000

1.659

-

!"# )

8.400

-

4.000

1.400

')* )

7.000

2.600

<

!"

)

)

Diameters aspect ratio ( ) 2.1.2 Experimental loop test

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The experimental loop test is shown schematically in Figure 3 and in a real view in Figure 4. The experimental facility is composed of gas and liquid circulation lines, tanks, booster pump, compressor, Variable Speed Drive (VSD), valves, measuring instrumentations and data acquisition system. The working fluids used are tap water and compressed air. The liquid phase flows in a closed loop, while the gas phase has an open circuit. These fluids do not follow the similarity established by Eq.(20), the reason for this is discussed in the next sections. Figure 3 illustrates the liquid line in green, which is composed of a centrifugal booster that pumps the liquid from the storage tank to the setup. The liquid flow rate is controlled by varying the rotational speed of the pump through the VSD. The air line is illustrated in blue in Figure 3, and is supplied by a compressor whose discharge is aligned to pulsation absorber tanks. A pressure regulator and needle valve control the air pressure and flow rate remotely. After the measurement of the gas and liquid flow rates through Coriolis Flow Meters, the phases are mixed at the setup inlet. A ring with four nozzles is used to inject the air into the liquid stream for the most homogeneous mixture possible at the inlet of the test section. Finally, the mixture runs through the capsule model, enters the intake and flows back to the liquid tank, according to the red line in Figure 3. In the tank, the gas is gravitationally separated and released, while the liquid returns to the line. In addition, in the red discharge line, a globe valve is installed in the return line, which is used to control the operating pressure by restricting the flow.

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Table 3. Dimensions of the real equipment and scaled model used in the experimental setup.

)

Hydraulic Diameter (

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2) Annular Duct: annular duct section upstream of the intake; 3) Gas-Liquid Separator: section of the annular duct just in front of the intake; 4) Undrained Top: dead zone at the top of the capsule where the separated gas accumulates. There are several diameter transitions, flanges and centralizers inside the capsule. The assumed dimensions are average and do not represent such details to simplify the construction of the test bench, which has unique diameters in the outer and internal ducts for simplicity. Moreover, these transitions and obstacles eliminate any possibility of the flow being developed within the capsule, so this poses no concern for making the experimental apparatus. Therefore, it is sufficient to scale the last section of the annular duct ( !F ) without variations in diameter or obstacles to the flow and the undrained top ( ')* ). Table 3 shows the average dimensions of the real equipment and the final dimensions of the test model. The reduced model was scaled from the similarity and performed with an aspect ratio of 2.72. This aspect ratio is based on the experimental limitations. The setup is composed by two 4m concentric tubes with diameters of 75 and 110 mm, resulting in a hydraulic diameter of 35 mm and diameter aspect ratio of 1.47. The outer tube is made of acrylic for flow visualization and the inner tube is made of PVC. The scale model does not represent the circular duct (LD) and annular duct 1 (LAN1) due to the nondevelopment of flow in this region, as previously discussed. Thus, Table 3 does not show these values.

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Figure 3. Layout of the experimental loop test.

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intake with pressure tap distances of 400 mm (∆ # ) and 560 mm (∆ ), and at the top of the annular duct (∆ ? ). For measuring the pressures ∆ # and ∆ , Validyne® transducers, model DP-15, were used. These sensors have an uncertainty of 0.25% of the full scale. The pressure ∆ ? were measured using a transducer manufactured by SMAR®, model LD-301, with uncertainty of 0.075% of the span. Before the phase mixture point, the liquid and gas flow rate were measured using Coriolis meters. For the liquid mass flow rate, two sensors were used according to the range. A 3-inch meter, series DS300 ( # ), and a 1-inch meter, series F100 ( ), were used for high and low flow rates, both manufactured by Emerson Micro Motion®. The gas flow rate measurement was performed by a 1/2-inch meter, series CMF015 ( ? ), from the same manufacturer. The operating temperature ( # ) is measured at the bench inlet by a PT-100 thermistor.

Figure 4. Experimental setup. 2.1.3 Instrumentation and Data Acquisition In the experimental setup, data acquisition and remote control of the equipment are performed by DAQ systems from National Instruments® and a LabVIEW® based software. The measuring instruments used and their locations in the setup are shown in Figure 3. The gauge pressures are measured at the setup inlet ( # ), upstream near the intake ( ), on the top of the annular duct ( ? ) and on the setup outlet ( > ). For measuring the gauge pressure, capacitive transducers, series 2088, manufactured by Rosemount® were used. The pressure sensors have an accuracy of 0.075%. For the differential pressure measurement, transducers were used at three different locations: upstream of the

2.1.4 Flow Visualization According to Mohammadi and Sharp (2013), the flow visualization using a high-speed camera offers the possibility to achieve high spatial and temporal resolutions and captures fast transient phenomena. In addition, highspeed imaging technique can be used to qualitatively identify and characterize flow patterns, and by processing the sequential images it is possible to acquire quantitative information about the flow. Therefore, to visualize the two-phase flow in the upstream annular duct and intake in the capsule, the technique of high-speed imaging was used.

Figure 5. Arrangement of high-speed cameras.

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The gas-liquid separation process is identified from the flow visualization, where we notice that not all the gas coming from the annular duct is sucked into the intake holes. The experimental procedure to calculate the separation efficiency, illustrated in Figure 6, consists of flooding the entire annular duct with liquid. This is accomplished by draining the gas eventually accumulated at the top of the capsule through the gas relief valve. The valve is then closed and gas injection ( ) starts at the capsule inlet. A fraction of the injected gas is dragged through the holes in the intake of the central duct ( , ) and another fraction is separated, accumulating at the top of the annular duct. Thereby, the liquid-gas interface will move downwards, as shown in Figure 6. The efficiency is then estimated by measuring the time required for the gas to occupy the top of the annular duct, until the gas-liquid interface reaches a standard reference mark, located at 1.98m from the top, i.e. the time required for the gas to fill a known volume. The process was performed 5 times for each experimental case of the test matrix. This procedure enables the average separation efficiency estimation. With time ( ) and measured flow rate ( ), it is estimated the injected gas volume ( ) and, from the position of the drained level, the separated gas volume at the top ( , ) is estimated and the average separation (Ē) can also be calculated:

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Two camera models were used: a Motion Pro X3 model camera, manufactured by RedLake®, which has a 1000 fps acquisition rate at a maximum resolution of 1280 x 1024, reaching 64000 fps at lower resolutions; and a Phantom VEO 640® model camera, which has a 1400 fps at maximum resolution of 2560 x 1600, reaching 360000 fps at lower resolutions. The cameras focused on two regions of interest: 1) the annular region, upstream the intake; 2) the intake region. For the annular region, the cameras can be positioned to provide a side or top view. For the region of intake, the camera can only be positioned laterally (side view). The cameras were attached to the capsule’s structure. Thus, the camera set inclines with the structure, i.e. has the same axis as the capsule. The cameras' positions are illustrated schematically in Figure 5. The illumination system consists of Xenon lamps and high-power LED reflectors. The data were acquired simultaneous to filming as the cameras’ triggers were synchronized with the data acquisition system through the LabVIEW® program.

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2.2 Experimental Procedure

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As previously mentioned, the experimental tests have two main objectives: 1) characterize the morphology of the gas-liquid two-phase within the capsule and 2) determine the separation efficiency. For the first objective, the experimental procedure consists of providing the desired phase flow rates, establishing the steady state and then performing the acquisition of data and images.

=

,

=

K

GL H

−

K GL

J

,

IJ

(21)

It is noteworthy that, as Eq. (21) dictates, this procedure obtains the average efficiency during the observation window and is therefore an estimation.

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Figure 6. Experimental procedure for estimating the average efficiency.

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The working fluids were tap water and compressed air. These fluid properties do not exactly follow the relationship defined in Eq. (20), resulting in incomplete similarity. However, the calculated properties of the liquid phase that follow the similarity are close to the water properties. For this reason, and for simplicity, in this first study, tap water was used as the liquid phase. The investigation of viscous effects, applied to other operational conditions, will be addressed in future studies. The experiments were conducted to cover as many points as possible regarding field reality. The real ESP Skid operates with liquid flow rates ranging from 500 to 5200 m3/d and no-slip void fractions between 10 and 50 %. Applying Eq. (15), the flow rates for testing the model are obtained, resulting in liquid superficial velocities ( # ) ranging from 0.1 to 1.0 m/s, and the same no-slip void fractions. These ranges of superficial velocity and void fraction were combined to provide an experimental matrix with 30 points, as shown in Table 4.

remaining 29 points were tested on the slopes of 5°, 15°, 30°, 45° and 60° to enable the detailed analysis of the gasliquid separation phenomenon, as well as the determination of flow patterns and parameters at the annular duct, and intake details. The test matrix points for the flow in the annular duct are plotted on the theoretical flow maps, as shown in Figure 7. The maps shown in Figure 7 were obtained from the model of Barnea et al. (1987) and add the modifications proposed by Caetano et al. (1992a, 1992b). In this figure, the abbreviations SS, SW, CH, SL, EB, B, AN and DB represent, the smooth stratified, wavy stratified, churn flow, slug flow, elongated bubbles, bubbles, annular, and dispersed bubbles flow patterns, respectively. Therefore, Figure 7 indicates that the expected flow pattern for all points on all experimental slopes will be the slug flow. It is important to highlight that the two-phase flow is not expected in this region. However, the observed characteristics of the flow are the ones felt by the equipment and that directly impact the flow segregation and efficiency. The Reynolds and Froude mixture numbers calculated for the test matrix indicate that the experiments covers laminar-to-turbulent regimes (1650 < < 30751) and subcritical-to-supercritical flow (0.18 < < 3.46).

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2.3 Experimental Matrix

QR 9T/V<

Table 4. Experimental Matrix with 30 experimental points.

10

20

0.10

1

0.19

MN%P 30

40

50

2

3

4

5

6

7

8

9

10

0.41

11

12

13

14

15

0.54

16

17

18

19

20

0.67

21

22

23

24

25

1.00

26

27

28

29

30

Point 1 was discarded due to its very low mass airflow, which made it difficult to maintain the steady state. The

of (b)

= 15°

(d)

= 45°

ro

= 5°

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(a)

= 30°

Jo (c)

(e)

= 60°

Figure 7. Theoretical flow pattern maps and test matrix points (black points) for different inclinations. The caption represents the flow patterns: SS - smooth stratified, SW - wavy stratified, CH - churn flow, SL - slug flow, EB - elongated bubbles, B - bubbles, AN - annular and DB - dispersed bubbles.

This section presents the results obtained in this experimental investigation. First, the analysis of the gasliquid flow morphology in the annular duct and intake region of the scaled capsule is presented. Then, the results and analysis about the separation efficiency are presented. 3.1 Flow Visualization in the annular duct

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Again, it is important to highlight the undeveloped characteristic of the flow inside the experimental setup. As there is almost no reference on such undeveloped flow, this discussion is important because the phenomena directly affect the separation efficiency and the flow in the intake region. Figure 8 shows the flow visualization of point 15, called P15, at 5° inclination, in the side view (a) and in the

top view (b); and Video 1 represents Figure 8b. This image shows the sequence of unit cells with the train of Taylor bubbles followed by a small amount of dispersed bubbles in the slug body. This is a classic initial arrangement for the slug flow pattern, but, unlike Caetano et al. (1992a,1992b) and Mendes (2012), it was also possible to observe that Taylor's bubbles did not involve the inner duct, remaining concentrated in the duct dorsum, which could be explained by the low inclination and aspect ratio of the duct. In general, the phase distribution in the 5° inclination is not axisymmetric, with the gas phase mostly at the duct’s dorsum, as shown in Figure 8a, which is expected for intermittent two-phase flow near horizontal position. The slug flow pattern is the most frequent in the test matrix for all inclinations.

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3 Results

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(a) side view

Figure 8. Flow in the annular duct for P15 9 # =

(b) top view

= 0.41 -/a; $ = 0.5<,

For higher liquid superficial velocities, bubble swarms were noticed, which made unit cell identification virtually impossible without a high-speed camera. This flow pattern was classified as the beginning of the transition from slug flow to dispersed bubbles. This is exemplified in Figure 9 and Video 2, which show the flow pattern of P28, at 5° inclination.

= 5° . Yellow arrow indicates flow direction.

For the same point in the experimental matrix, the increase in the slope of the installation promotes an increase in the irregularity of the bubbles and unit cells, an increase in the dispersed bubbles in the slug bodies, and a decrease in the elongated bubble size (Figure 10).

Figure 9. Flow in the annular duct for P28 9 # = 1,0 -/a; = 0.43 -/a; $ = 0.3<, = 5° (side view). Yellow arrow indicates flow direction and red arrow indicates an elongated bubble.

This can be seen in Figure 10, which represents P15, for inclinations ranging from 15 to 60°. Figure 10a clearly presents large bubbles and for the same operational condition; as the slope increases, the bubble swarm increases (Figure 10b), but larger elongated bubbles are

still identified. Figure 10c and Figure 10d show that the bubbles become irregular and smaller in size leading to an initial transition to bubble flow. However, as the unit cells are still observed, these points are classified as slug flow.

= 15°

(b)

= 30°

(c)

= 45°

(d)

= 60°

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(a)

Figure 10. Flow pattern characteristics in the annular duct with the increase in the slope for P15 9 0.5<, in top view. Yellow arrow indicates flow direction. As in the 5° slope, the 15° and 30° slopes showed slug flow patterns for almost all cases of the test matrix (Figure 11a and Figure 12a). However, as the liquid and gas velocities increases, the elongated bubbles decreases, which also decrease in size, change in shape and an increase in bubble swarm in the slug body is noticed (Figure 11b and Figure 12b). As the slope increases to 45° and 60°, the void fraction increases the irregular bubbles, which in turn causes the increase in dispersed bubbles in the slug body, it is accentuated in the annular duct before the intake region (Figure 13 and Figure 14). In the 45° and 60° slopes, the

#

=

= 0.41 -/a; $ =

flow patterns are less defined. Both slopes show a change of flow pattern from slug to dispersed bubbles for the cases of lower gas velocity and high liquid velocity (Figure 14b). However, were there to be a detailed description, in some cases, a transition from slug flow to churn (Figure 13b), and bubbles to slug flow (Figure 13a and Figure 14b) could be speculated. There is also a discussion about the existence of the cap bubbly flow pattern, as in Sun et al. (2004), and some of the observed flow patterns, mainly those with large bubbles, but smaller than the Taylor bubbles, which might be classified as cap-bubbly. However, this is not discussed

Table 5 highlights all the flow patterns seen for each case on each slope. As already discussed, only the slug flow and dispersed bubble flow patterns were clearly identified. The points in which the dispersed bubble flow were observed are marked with a black ellipse in Figure 7, for 45 and 60° degrees. The results of this observation are used for the analysis of the separation process data (section 3.3). As the experimental bench inclination increased, in addition to the bubbles becoming more irregular and dispersed, as already discussed, we observed that for the lower liquid superficial velocities (J1 = 0.10, 0.19, 0.41 and 0.54 m/s), the bubbles move from side to side as they ascend to the dorsum of the annulus upstream of the intake. In some cases, larger zigzag blisters were identified in the aqueous medium (Figure 15 – yellow arrows, P03 of the test matrix), as if they could surround the inner duct. Additionally, Video 3 illustrates Figure 15a. These phenomena were observed more frequently and at the lowest mixture superficial velocity points for slopes of 45° and 60°, this phenomenon occurred throughout the test matrix.

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in this article due to the non-conclusive work regarding this classification. Finally, as with the present case is of undeveloped flows, only slug flow, dispersed bubbles and transitions between them are clearly reported here. When the unit cell is observed, even with small or distorted bubbles, the flow is classified as slug flow (SL), and when only dispersed bubbles are observed, the flow is classified as dispersed bubbles (DB). A more detailed analysis of flow patterns is possible but would be speculative. Based on what was observed in the images recorded by the high-speed cameras,

(a) P10 –

(b) P30 –

#

#

= 0.19 -/a,

= 1.00 -/a,

= 0.19 -/a, $ = 0.5 – Slug Flow

= 1.00 -/a, $ = 0.5 – Slug/Dispersed Bubble Flow

Figure 11. Observed flow patterns at 15° (top view). Yellow arrow indicates flow direction.

(a) P18 –

#

= 0.54 -/a,

= 0.23 -/a, $ = 0.3 – Slug Flow

#

(b) P29 –

= 1.00 -/a,

= 0.67 -/a, $ = 0.4 – Slug/Dispersed Flow

#

= 0.19 -/a,

= 0.08 -/a, $ = 0.3 – Slug Flow

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(a) P08 –

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Figure 12. Observed flow patterns at 30° slope (top view). Yellow arrow indicates flow direction.

#

= 0.41 -/a,

= 0.41 -/a, $ = 0.5 – Slug Flow

#

= 0.67 -/a,

= 0.17 -/a, $ = 0.2 – Slug Flow

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(b) P15 –

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Figure 13. Observed flow patterns in 45° slope (top view). Yellow arrow indicates flow direction.

(a) P22 –

(b) P26 –

#

= 1.00 -/a,

= 0.11 -/a, $ = 0.1 – Dispersed Bubbles

Figure 14. Observed flow patterns in 60° slope (top view). Yellow arrow indicates flow direction.

Table 5. Classification of flow patterns for the test matrix. 15° SL SL SL SL SL SL SL SL SL SL SL/DB SL/DB SL/DB SL/DB SL/DB

(a) P03 – 15° –

(b) P18 – 30° –

#

#

30° SL SL SL SL SL SL SL SL SL SL SL/DB SL/DB SL/DB SL/DB SL/DB

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5° SL SL SL SL SL SL SL SL SL SL SL/DB SL/DB SL/DB SL/DB SL/DB

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Point 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30

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60° SL SL SL SL SL SL SL SL SL SL SL SL SL SL

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45° SL SL SL SL SL SL SL SL SL SL SL SL SL SL

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30° SL SL SL SL SL SL SL SL SL SL SL SL SL SL

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15° SL SL SL SL SL SL SL SL SL SL SL SL SL SL

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5° SL SL SL SL SL SL SL SL SL SL SL SL SL SL

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Point 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15

= 0.10 -/a,

= 0.04 -/a, $ = 0.3

= 0.54 -/a,

= 0.23 -/a, $ = 0.3

45° SL SL SL SL SL DB SL SL SL SL DB DB SL/DB SL/DB SL/DB

60° SL SL SL SL SL DB SL SL SL SL DB DB SL/DB SL/DB SL/DB

#

= 0.19 -/a,

= 0.19 -/a, $ = 0.5

(d) P16 – 60° –

#

= 0.54 -/a,

= 0.06 -/a, $ = 0.1

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(c) P10 – 45° –

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Figure 15. Zigzag movement of bubbles in the annular duct (top view). The flow is from the lhs to the rhs.

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3.2 Flow Visualization in the intake

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In the intake region, part of the air bubbles were dredged by the intake holes and part rose straight to the annular top section for all inclinations, as shown in Figure 16 for P04 and in Figure 17 for P13, both at 5, 45° and 60°. Video 4 and Video 5 refer to Figure 16a and Figure 16c, respectively; and Video 6 refers to Figure 17b. The oscillation of the dynamic liquid-air level was also noticed, according to the increase of the slope (Figure 16b and Figure 16 for P04); this change in level is more evident in the visualization of the cases of lower liquid flow.

(a)

= 5°

(b)

= 45°

(c)

= 60°

Figure 16. Flow in the intake region for P04, # = 0.10 -/a, = 0.06 -/a, $ = 0.4 (side view). The yellow arrows indicate the flow direction and the yellow dashed line is the dynamic gas-liquid level. The yellow dashed line is the dynamic gas-liquid level in (b) and (c). The flow is from the lhs to the rhs. Also recurrently observed, in all slopes, were secondary flows in the intake region, such as bathtub type vortices in the holes, caused by pressure differences in the region. An example is shown in Figure 18 and for P28 in different slopes. Video 7 represents Figure 18c. Although these vortices are present in all slopes, they were more frequently observed in the 45° and 60° inclinations and grew in intensity as the mixture velocity increased.

(a)

= 5°

= 45°

(c)

= 60°

(b) 45°

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(c) 60°

Figure 18. Secondary flows and bathtub vortices indicated by yellow/red arrows. Flow in the intake region for P28, = 0.43 -/a, $ = 0.3 (side view). The # = 1.00 -/a, flow is from the lhs to the rhs.

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Figure 17. Flow in the intake region for P13, # = 0.41 -/a, = 0.18 -/a, $ = 0.3 (side view). The yellow arrows indicate the flow direction. The flow is from the lhs to the rhs.

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The presence of secondary flows in the intake, as reported, reinforces the need to evaluate their impact on the separation process from the use of CFD techniques or PIV measurements. 3.3 Separation efficiency

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As observed in Figure 16, Figure 17 and Figure 18 the gas-liquid separation process is clear from the flow visualization, where it is clear that not all the gas coming from the annular duct is sucked into the intake holes. During the experiments, the average separation efficiency was estimated for each case of the experimental matrix, as described in Section 2.2 and Figure 6. The average separation efficiencies were obtained for the test matrix previously defined in Table 4, where it is presented as a function of the mixture velocity (Jm) and the no-slip void fraction (λ), ranging from 10 to 50%. The experiments were performed at 5°, 15°, 30°, 45° and 60°, and the results are shown in Figure 19.

(a) 5°

The Froude number is defined as the relation between inertial forces and another exerted by a field (gravity or rotation), important in segregated flows, as defined in Eq. (10). The only variable for each inclination in the Froude number is the mixture flow rate, or the mixture velocity, so the graphs are shown with the mixture velocity in the xaxis. As observed in Figure 19, in general, the separation efficiency decreases as the mixture velocity increases, which is expected. However, there are some operational conditions for 45° and 60° where the trend changes and the efficiency starts to increase (dark points inside red circles). In this case, changes are observed in the flow pattern from slug flow to dispersed bubbles. This fact will be the subject of analysis in the sequence of this article. The uncertainty of the separation efficiency was calculated based on the standard deviation of 5 measurements for each point. In most cases the uncertainty was lower than 1%. It was observed that for the lowest liquid velocities (0.1 m/s) the standard deviation increases, but it stays lower than 4%. The error bars were also show in Figure 19. Alhanati (1993) analyzed the same efficiency as a function of the mixture velocities and proposed the model shown in Eq. (22): = 1−f

1−

gh

1−$ i $

(22)

where α is the gas void fraction calculated through Eq. (23) from the drift-flux model:

uv [-]

wkxy [m/s]

15°

1.2

0.72

30°

1.2

0.76

45°

1.2

0.75

60°

1.2

0.69

Inclination [°]

1.2

5°

The model proposed by Bendiksen (1984) was adapted to calculate Taylor's bubble velocity 9 k'l <, incorporating the revised models for intermittent flow in annular ducts, as shown in Eq.(24). In this case, the model from Bendiksen (1984) was written for the two-phase Froude number using the external diameter (De) on the aspect ratio, as suggested by Hasan and Kabir (1992) and proven by Mendes (2012). For the distribution coefficient, C0=1.2 was used, as suggested for intermittent flow. k'l

= N0.5491 + 0.185
0.34591 + 0.29
∆qrst

(24)

qr

re

Table 6 shows the distribution coefficient and Taylor’s bubble velocity calculated for all cases.

The separation efficiency calculated with the model proposed by Alhanati (1993) is shown in Figure 19, compared to the results obtained experimentally. The results expressed in Figure 19 are qualitatively in line with what was expected, where efficiencies tend to be 100% when liquid flow tends to zero (no bubbles would be dragged through the intake and all gas would go to the top) and when the liquid flow tends to infinity (all bubbles are sucked in due to the drag of the liquid phase), as expected and reported in the literature (Alhanati, 1993; Alhanati et al., 1994; Lea and Bearden, 1982b).

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jL +

(23)

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=

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Table 6. Drift-flux model parameter: Distribution coefficient and Taylor’s bubble velocity values.

(a)

= 5°

0.67

(b)

= 15°

of

= 30°

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(d)

= 45°

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(c)

(e)

= 60°

Figure 19. Separation Efficiency as a function of mixture velocity for different gas fractions. Dark square points inside the red regions are the ones with different trends. However, when comparing the magnitude of the results with those estimated to the model proposed by Alhanati (1993), large deviations are observed, as illustrated in Figure 19. One of the points is the calculus of the void fraction with drift flux-model parameters that are not specific for this case. In addition, the model proposed by Alhanati (1993) disregards the radial slippage. This is another point that affect the model prediction compared to the experimental results. Finally, the model was proposed for vertical flow, unlike the experiments performed in this article. These points will be discussed in further articles devoted to the modeling of the observed phenomena. For the 45° and 60° slopes, inflection points of separation efficiency (black square symbols within the red regions in the graphs) were observed in the cases of lower no-slip void fraction (10% and 20%) with larger mixture velocities (Jm). At these points, changes in the flow pattern

from slug flow to dispersed bubbles were identified, given in the images in Figure 20 (a) and Video 8 for point P26 at 45°, Figure 20 (b) and (c) for points P21 and P27 at 60°. It was possible to conclude that this change in the flow pattern, from intermittent to dispersed flow patterns, modifies the buoyancy and drag forces acting in the gas phase changes, thus altering the separation efficiency. Also, curves were plotted for the separation efficiency as a function of the mixture velocity for the same no-slip void fraction for different slopes to visualize any possible disparity between them, as shown in Figure 21. After analyzing the graphs, it was possible to verify that the separation efficiencies are not significantly influenced by the slope. Except for the cases with the lowest void fractions (10% and 20%) with higher mixture velocities at 45° and 60° slopes, where, as already

discussed, changes in the flow pattern from slug flow to dispersed bubbles were observed.

The results observed here were used for mathematical modeling of the gas-liquid two-phase flow separation, which will be the scope of further articles.

#

= 1.00 -/a,

= 0.11 -/a, $ = 0.1 – Dispersed Bubbles

(b) P21 – 60° –

#

= 0.67 -/a,

= 0.07 -/a, $ = 0.1 – Dispersed Bubbles

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(a) P26 – 45° –

#

= 1.00 -/a,

ur

(c) P27 – 60° –

= 0.25 -/a, $ = 0.2 – Dispersed Bubbles

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Figure 20. Cases where dispersed bubble flow patterns where observed and there is inflexion in efficiency separation.

(a) $ = 10%

(b) $ = 20%

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(d) $ = 40%

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(c) $ = 30%

9 < $ = 50%

Figure 21. Separation Efficiency for the same no-slip gas fraction at different inclinations as a function of mixture velocity.

4 Conclusions This article presents new experimental results and observations on two-phase gas-liquid flows in annular ducts. Flow patterns in the annulus of such undeveloped flow were observed and discussed. Also, the flow in the region of the intake (holes in the experimental setup) was observed as was the presence of a dynamic level of separation. Finally, an analysis of the separation efficiency was carried out, where the data was compared with a model from the literature. The main conclusions are highlighted below: • The model by Barnea et al. (1987), added to the modifications proposed by Caetano et al. (1992) for vertical flow in annular ducts, correctly predicted that all points of the test matrix would operate in the slug flow pattern, observed at 5°, however, as the slope

increases, the flow pattern transitions to dispersed bubbles, which indicates the necessity of amelioration on such models. However, it is important to highlight that this is a case of undeveloped flow and more studies about flow patterns in such condition are required • Several details of flow in annular ducts were reported, such as the presence of bubble zigzag as the pipe inclination increases.

• The intake region where the gas phase is sucked was detailed by observing bathtub vortices and differences in drag at different mixture velocities. Still, the images show the detailed movement of bubbles in this region. • As observed in the literature, there is an exponential decay of separation efficiency as the mixture velocity increases, which compares to increasing the Froude number. This shows that, when inertial effects become greater than those of buoyancy, the liquid phase carries

Caetano, E. F., Shoham, O., Brill, J.P., 1992b. Upward Vertical Two-Phase Flow Through an Annulus - Part II: Modeling Bubble, Slug, and Annular Flow. Journal of Energy Resource Technology 114, 14-30. doi:10.1115/1.2905916 Colodette, G., Pereira, C.A., Siqueira, C.A.M, Ribeiro, M.P., 2007. The New Deep Water Oil and Gas Province in Brazil: Flow Assurance and Artificial Lift: Innovations for Jubarte Heavy Oil. Society of Petroleum Engineers (SPE). 9p. doi:10.4043/19083-MS

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Costa, B.M.P., Oliveira, P. da S., Roberto, M.A.R., 2013. Mudline ESP: Electrical Submersible Pump Installed in A Subsea Skid. Society of Petroleum Engineers (SPE). 10p. doi:10.4043/24201-MS

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Estevam, V., 2002. A Phenomenological Analisys About Centrifugal Pump in Two-Phase Flow Operation. 265p, PhD Thesis, University of Campinas. Campinas. Harun, A.F., Prado, M.G., Serrano, J.V., Doty, D.R., 2001. A Mechanistic Model to Predict Natural Gas Separation Efficiency in Inclined Pumping Wells. SPE Production Operations Symposium. 9p. https://doi.org/10.2118/67184-MS

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practically all the gas through the intake. This is observed for all slopes in all no-slip void fractions. • The literature model (Alhanati, 1993) was unable to correctly predict the separation process. However, it did predict the same trend. This could be due to the void fraction prediction by the drift-flux model with the parameters not adapted to the present case and the disregard for the slippage between phases in the radial direction. These points will be discussed in further articles devoted to the modeling of the observed phenomena. • In some cases of low no-slip void fraction and high mixture velocities, the separation trend changes. Through the experimental observations, changes in the flow patterns from slug to dispersed flow were responsible for those changes. We hope the results of this article will be valuable for further studies on CFD and modeling of such flows. Also, these results will be used for the modeling of gas-liquid separation analysis and comparison with other literature models in further articles.

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5 Acknowledgements

6 References

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The authors would like to thank Petrobras (Petróleo Brasileiro S/A) and ANP (“Compromisso de Investimentos com Pesquisa e Desenvolvimento”) for providing financial support for this study. The authors also thank the Artificial Lift & Flow Assurance Research Group (ALFA), the Center for Petroleum Studies (CEPETRO), and the School of Mechanical Engineering (FEM), all part of the University of Campinas (UNICAMP).

Alhanati, F.J.S., 1993. Bottomhole Gas Separation Efficiency in Electrical Submersible Pump Installations. 280p. PhD. Thesis, University of Tulsa, Tulsa.

Hasan, A.R., Kabir, C.S., 1992. Two-Phase Flow in Vertical and Inclined Annuli. International journal of Multiphase Flow 18, 279-293. doi:10.1016/03019322(92)90089-Y Lackner, G., 1997. The Effect of Viscosity on Downhole Gas Separation in a Rotary Gas Separator. PhD. Dissertation. University of Tulsa, Tulsa. Lea, J.F., Bearden, J.L., 1982. Effect of Gaseous Fluids on Submersible Pump Performance. Journal of Petroleum Technology 34, 2922-2930. doi:10.2118/9218-PA

Lea, J.F., Bearden, J.L., 1982a. Gas Separator Performance for Submersible Pump Operation. Journal of Petroleum Technology 34, 1327-1333. doi:10.2118/9219-PA

Barnea, D., 1987. A Unified Model for Predicting FlowPattern Transitions for the Whole Range of Pipe Inclinations. International Journal of Multiphase Flow 13, 1-12. doi:/10.1016/0301-9322(87)90002-4

Liu, B., 2002. Modeling Down-hole Natural Separation Using Bubble Tracking Method. MSc. Dissertation. University of Tulsa, Tulsa.

Bendiksen, K.H., 1984. An experimental investigation of the motion of long bubbles in inclined tubes. International Journal of Multiphase Flow 10, 467-483. doi:10.1016/0301-9322(84)90057-0

Liu, B., Prado, M., 2004. Application of a Bubble Tracking Technique for Estimating Downhole Natural Separation Efficiency. Journal of Canadian Petroleum Technology, 43(05), 57-61. doi:10.2118/04-05-05

Caetano, E. F., Shoham, O., Brill, J.P., 1992a. Upward Vertical Two-Phase Flow Through an Annulus - Part I: Single-Phase Friction Factor, Taylor Bubble Rise Velocity, and Flow Pattern Prediction. Journal of Energy Resource Technology 114, 1-13. doi:10.1115/1.2905917

Marquez, R. Prado, M., 2001. Consolidation of The Experimental Data for Natural Separation Efficiency. The University of Tulsa, Artificial Lift ProjectsTUALP. Technical Report TR-07.

Mohammadi, M., Sharp, K.V., 2013. Experimental techniques for bubble dynamics analysis in microchannels: A review. Journal of Fluids Engineering, Transactions of the ASME 135(2), 10p. doi:10.1115/1.4023450

Teixeira, V.F., Gessner, T.R., Shigueoka, I.T. 2012. Transient Modeling of a Subsea Pumping Module Using an ESP. SPE Latin American and Caribbean Petroleum Engineering Conference, 12p. doi:10.2118/153140-MS Vieira, S.C., 2018. Gas-Liquid Separation on the Intermittent Two-Phase Flow in an Annular Duct. 233p, MSc Dissertation, University of Campinas, Campinas.

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Monte Verde, W., Biazussi, J.L., Sassim, N. A., Bannwart, A.C., 2017. Experimental Study of Gas-Liquid TwoPhase Flow Patterns Within Centrifugal Pumps Impellers. Experimental Thermal and Fluid Science 85, 37-51. doi:10.1016/j.expthermflusci.2017.02.019

Tarcha, B.A., Furtado, R.G., Borges, O.C., Vergara, L., Watson, A.I., Harris, G.T., 2016. Subsea ESP Skid Production System for Jubarte Field. Offshore Technology Conference. Offshore Technology Conference, 20p. doi: 10.4043/27138-MS

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Mendes, F.A.A., 2012. Experimental Study, numerical simulation and phenomenological modeling of gravitational separation of gas in down-hole directional wells. 250p. PhD Thesis, University of São Paulo - São Carlos School of Engineering. São Carlos.

Tarcha, B.A., Borges, O.C., Furtado, R.G., 2015. ESP Installed in a Subsea Skid at Jubarte Field. SPE Artificial Lift Conference - Latin America and Caribbean, 8p. doi: 10.2118/173931-MS

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Marquez, R., Prado, M., 2003. A New Robust Model for Natural Separation Efficiency. SPE Production Operations Symposium, 11p. doi:10.2118/80922-MS

Taitel, Y., Barnea, D., Dukler, A., 1980. Modelling Flow Pattern Transitions for Steady Upward Gas-Liquid Flow in Vertical Tubes. AIChE Journal. 26(3), 345354. doi:10.1002/aic.690260304

-p

Marquez, R., 2004. Modeling Downhole Natural Separation. PhD. Thesis. The University of Tulsa, Tulsa.

Panton, R.L., 2013. Incompressible Flow, 4th ed. Wiley, New Jersey. doi:10.1002/9781118713075

ur

na

Rodrigues, R., Soares, R., Matos, J.S.D., Pereira, C. a G., Ribeiro, G.S., 2005. A New Approach for Subsea Boosting - Pumping Module On the Seabed. Offshore Technology Conference, 8p. doi:10.4043/17398-MS.

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Sambangi, S.R., 1994. Gas Separation Efficiency in Electrical Submersible Pump Installations with Rotary Gas Separation. 70p. MSc. Dissertation. University of Tulsa, Tulsa. Shoham, O., 2005. Mechanistic Modeling of Gas-Liquid Two-Phase Flow in Pipes, 408p. Society of Petroleum Engineers. Sun, X., Kuran, S., Ishii, M., 2004. Cap bubbly-to-slug flow regime transition in a vertical annulus. Experiments in Fluids 37, 458-464. doi:10.1007/s00348-004-0809-z

Vieira, S.C., Geest, C. der, Todorovic, A.F., Castro, M.S. De, Bannwart, A., 2020. Experimental Study and Modelling of Developing Intermittent Slug Flow in Inclined Annular Pipelines. Journal of Petroleum Science and Engineering, 192, 107204. doi:10.1016/j.petrol.2020.107204 Viloria, J.C.S., 1999. Natural Separation Efficiency in Electric Submersible Pump Systems. 83p. MSc. Dissertation. University of Tulsa, Tulsa. Wallis, G.B., 1969. One-dimensional two-phase flow. McGraw‐Hill, 408p. New York.

White, F.M., 1998. Fluid mechanics, 4th ed., 944p, McGraw-Hill Book Company. White, F.M., 2006. Viscous Fluid Flow, 2nd ed. 629p. McGraw-Hill.

Highlights

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Experimental investigation of the gas-liquid separation efficiency in a scaled ESP skid. Undeveloped flow patterns inside annular ducts are described. Gas-liquid separation efficiency is a function of flow pattern, mixture velocity, no-slip void fraction, slope and mixture Froude number. Trend in separation efficiency is changed due to flow pattern transition from slug to dispersed bubbles due to buoyancy and drag forces change. Secondary flows and bathtub vortices are present in the intake.

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• • •

Declaration of interest

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The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.