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T H E U N I V E R S I T Y O F T U L S A

THE GRADUATE SCHOOL

DESIGN AND PERFORMANCE OF MULTIPHASE

DISTRIBUTION MANIFOLD

by

Angel R. Bustamante

A thesis submitted in partial fulfillment of

the requirements for the degree of Master of Science

in the Discipline of Mechanical Engineering

The Graduate School

The University of Tulsa

2003

iii

ABSTRACT

Bustamante, Angel R. (Master of Science in Mechanical Engineering)

Design and Performance of Multiphase Distribution Manifold (76 pp.- Chapter VI)

Directed by Dr. Ram Mohan, Dr. Shoubo Wang and Dr. Ovadia Shoham.

(150 words)

A novel multiphase distribution manifold is studied experimentally and

theoretically. The distribution manifold is elevated, with slug dampers and GLCCs

attached downstream, as an integrated system. The objective of the novel manifold

system is to gather production from several inlet wells, and to ensure equal split of gas

and liquid phases for downstream equipment.

Over 200 experimental runs were conducted, evaluating eight different inlet well

configurations. Individual inlet well gas and liquid flow rates and liquid and gas split

ratios were measured. Data analysis revealed that the capability of the distribution

manifold system is to ensure fairly equal split ratios downstream, for different inlet flow

conditions.

A mechanistic model is developed, based on the Hardy-Cross method, capable of

predicting the downstream liquid and gas split ratios. Good agreement is observed

between the model predictions and the experimental data for all tested inlet flow

conditions, showing errors between 3% to 15%.

iv

ACKNOWLEDGMENTS

I am extremely grateful to Dr. Ram Mohan for his unconditional support; to Dr.

Ovadia Shoham for his continous help at any time; to Dr. Shoubo Wang and Dr. Luis

Gomez for their support throughout this investigation, and, more than that, for being my

friends. I also wish to thank Dr. Brenton McLaury, member of the Thesis Committee, for

his suggestions and comments. I am also indebted to Dr. Gene Kouba from

ChevronTexaco, for initiating this project and for serving on my Thesis Committee.

I want to thank the following persons and entities for their support and guidance

during my study and research:

• PDVSA, for giving me the chance to come to TU and support me, enabling

me to accomplish this achievement.

• Mr. Carlos Torres, my office mate and friend, for his help throughout my

research, especially while developing the mechanistic model.

• Ms. Judy Teal for her help, support and encouragement throughout my stay

at TU.

• TUSTP graduate students and member companies for their valuable

comments, cooperation and friendship during this project.

• The U.S. Department of Energy (DOE) for supporting this project.

• To all my friends who share with me this important period of my life.

v

DEDICATION

This work is dedicated to my lovely wife Nivia for her support, encouragement

and love. I would also like to dedicate this work to my parents Miguel and Juana, my

family, and especially my youngest brother Miguel.

vi

TABLE OF CONTENTS

APPROVAL PAGE ii

ABSTRACT iii

ACKNOWLEDGMENTS iv

DEDICATION v

TABLE OF CONTENTS vi

LIST OF FIGURES viii

I. INTRODUCTION 1

II. LITERATURE REVIEW 5

2.1. Splitting at Impacting Tee 5

2.2. Two-Phase Flow in T-Junctions 6

2.3. Slug Flow 7

2.4. Manifolds 8

2.5. Flow Conditioning 8

III. EXPERIMENTAL PROGRAM 10

3.1. Experimental Facility 10

3.1.1. Metering and Storage Sections 11

3.1.2. Modular Test Section 12

3.1.3. Instrumentation, Control and Data Acquisition System 13

3.1.4. Multiphase Distribution Manifold/Slug Damper

/GLCC System 14

3.2. Experimental Setup 19

vii

3.2.1. Parameter Definitions 20

3.2.2. Test Procedure 22

3.2.3. Test Matrix 22

3.3. Experimental Results. 24

3.3.1. Liquid Carry-Over Operational Envelope of Integrated System 24

3.3.2. Manifold Operational Envelope for Liquid Carry-Over 27

3.3.3. Liquid and Gas Split Ratios in Distribution Manifold 29

3.3.4. Resistance Coefficient of Manifold 32

3.3.5. Transient Performance of Multiphase Distribution Manifold 36

IV. MODEL DEVELOPMENT 40

4.1. Manifold Design Model 40

4.1.1. Manifold Sizing 40

4.1.2. Gas and Liquid Outlets sizing 46

4.2. Manifold / Slug Damper / GLCC System Performance Model 49

4.2.1. Calculation of liquid height at node “a” 53

4.2.2. Calculation of Head Losses between nodes 54

V. SIMULATIONS AND RESULTS 59

5.1. Liquid Split Ratio 59

5.2. Gas Split Ratio 64

VI. CONCLUSIONS AND RECOMMENDATIONS 68

NOMENCLATURE 72

REFERENCES 74

viii

LIST OF FIGURES

Figure 1.1. GLCC Separator 2

Figure 1.2. Multiphase Distribution Manifold Schematic 3

Figure 3.1. Schematic of Experimental Facility 10

Figure 3.2. Tanks, Pumping Station and Metering Section 12

Figure 3.3. Schematic of Manifold / Slug Damper / GLCC System 15

Figure 3.4. Liquid and Gas Inlet Flow Meters 16

Figure 3.5. Multiphase Distribution Manifold Schematic 17

Figure 3.6. Slug Damper Manifold 18

Figure 3.7. GLCC Separator Schematic 19

Figure 3.8. Flow Configurations in Inlet Wells 23

Figure 3.9. Operational Envelope for LCO of Manifold / Slug Damper /

GLCC System 26

Figure 3.10. Operational Envelope of the Distribution Manifold 28

Figure 3.11. Liquid Split Ratio in the Distribution Manifold 30

Figure 3.12. Gas Split Ratio in the Distribution Manifold 31

Figure 3.13. Resistance Coefficient of Liquid Outlet vs. Reynolds Number 34

Figure 3.14. Conditions for Determination of Resistance Coefficient 35

Figure 3.15. Transient Performance of Manifold Under No Gas Flow 37

Figure 3.16. Transient Performance of Manifold Under Low Gas Flow 38

Figure 3.17. Transient Performance of Manifold Under Moderate Gas Flow 39

Figure 4.1. Stationary Finite Wave on Gas-Liquid Interface 41

ix

Figure 4.2. Manifold Sizing-Criterion 1 44

Figure 4.3. Manifold Sizing-Criterion 2 45

Figure 4.4. Liquid and Gas Split Calculations 45

Figure 4.5. Variables Involved in Liquid Outlet Sizing 46

Figure 4.6. Variables Involved in Manifold Performance Model 50

Figure 4.7. Simplified Piping and Nodes Model 52

Figure 5.1. Comparison between Model Predictions and Experimental Data for

Liquid Split Ratio in Flow Configuration I 60


Liquid Split Ratio in Flow Configuration III 61


Liquid Split Ratio in Flow Configuration VI 62


Liquid Split Ratio in Flow Configuration IV 63


Gas Split Ratio in Flow Configuration I 64


Gas Split Ratio in Flow Configuration III 65


Gas Split Ratio in Flow Configuration VI 66


Gas Split Ratio in Flow Configuration IV 67

CHAPTER I

INTRODUCTION

During the past decade the petroleum industry has been forced to seek less

expensive and more efficient alternatives to conventional gravity-based separators.

Compact separation systems represent a key factor in the reduction of costs of oil and gas

production and handling. The so-called Gas-Liquid Cylindrical Cyclone (GLCC©1) is a

successful example of how a simple idea can provide a solution to the above-mentioned

need of the petroleum industry. The GLCC is a compact, cheap and lightweight separator

that requires almost no maintenance, is easy to operate and construct and has a small

footprint. Thus, the GLCC is an economically attractive alternative to the big, heavy and

expensive conventional gravity-based separator.

The GLCC, shown schematically in Figure 1.1, is simply a vertical piece of pipe

with a downward inclined tangential inlet and two outlets, one at the top for gas and one

at the bottom for liquid. An important feature of the GLCC is that it does not have

moving parts or any internal devices, which make it easy to operate and maintain. The

inclined tangential inlet provides a swirling motion in the GLCC, whereby due to

gravitational and centrifugal forces the gas and liquid phases are separated. The liquid is

forced towards the wall and leaves the GLCC from the bottom outlet, whereas the gas

moves to the center of the cylinder and exits from the top.

One of the problems with the operation of GLCC is its low residence time. This

feature causes operational problems when flow fluctuations occur. To solve this problem,

1 GLCC© Gas-Liquid Cylindrical Cyclone-copyright, The University of Tulsa, 1994

2

the GLCC can be equipped with a robust control system that maintains a constant liquid

level inside the separator. However, this solution causes problems downstream of the

GLCC because large outlet liquid flow rate fluctuations may occur during the production

of large slugs, much higher than the average liquid flow rate. To improve the

performance of the separation equipment some inlet flow conditioning devices can

provide an effective and cheap solution by smoothening the flow fluctuations. Reinoso

(2002) analyzed one such device, namely the “Slug Damper”. This device not only

protects downstream equipment but also extends the operational envelope of GLCC.

When a terrain slug hits the damper, the larger slug is dampened, providing fairly

constant flow rate into the downstream GLCC.

This study presents the Multiphase Distribution Manifold, which can be used in

conjunction with the slug damper and the GLCC. This integrated system minimizes

downstream operational problems, by dampening large flow variations and distributing

the liquid and gas flow equally. The proposed distribution manifold, shown schematically

Figure 1.1. GLCC Separator

3

in Figure 1.2, is a simple horizontal pipe section, into which different wells are

connected. As can be seen, upper gas outlets and lower liquid outlets are provided. This

novel device can function as a pre-separator upstream of compact separators. Also, the

distribution manifold attempts to provide equal liquid and gas split for downstream

processing facilities, regardless of the variation of flow pattern and flow rates in the

different inlet wells.

Another important feature of the proposed distribution manifold, when working in

conjunction with a slug damper, is the increment of the damping capacity of the system,

due to the available additional volume to dampen large slugs before entering the GLCCs.

Research goals and objectives: The main goal of this study is to acquire

experimental data in order to determine the liquid and gas split ratios in the distribution

manifold system for different flow conditions in each inlet well and different flow

Figure 1.2. Multiphase Distribution Manifold Schematic

Distribution Manifold

Inlet Wells

Liquid Outlets

GasOutlets


Inlet Wells

Liquid Outlets

GasOutlets

4

configurations inside the manifold. This study also intends to identify the mechanisms

involved in the distribution manifold performance, so as to enable the development of a

mechanistic model for design and performance prediction purposes.

The specific objectives of this study are as follows:

• Design, construct and install a multiphase distribution manifold facility.

• Instrument the multiphase distribution manifold facility to enable

measurement and monitoring of the outlet liquid and gas flow split.

• Acquire experimental data of liquid and gas split ratios for different flow rates

in each inlet well and different flow configurations in the manifold.

• Identify and characterize the mechanisms involved in the liquid and gas split

process in the distribution manifold to shed light on the complex

hydrodynamic behavior of the distribution manifold.

• Develop a mechanistic model as a tool to design a distribution manifold that

aims at equal liquid and gas flow rates in each downstream separator. The

mechanistic model can also be used to evaluate the performance of the

distribution manifold when its geometry is given.

The next chapter encompasses a review of the literature relevant to this study. In

Chapter III, the experimental investigation is presented, which includes description of the

facilities, experimental program and experimental results. Chapter IV presents the

developed mechanistic model, and Chapter V shows a comparative study between the

model predictions and the experimental data. Conclusions of this investigation are

summarized in Chapter VI along with some recommendations for future work. This is

followed by a list of nomenclature and references in separate sections.

CHAPTER II

LITERATURE REVIEW

The proposed Multiphase Distribution Manifold system configuration is a

relatively new device for flow conditioning upstream of compact separators. There are

few literature references related to the performance of individual components of the

manifold system, but almost none related to the design and evaluation of the distribution

manifold as a whole. This chapter presents an overview of pertinent literature studies,

such as two-phase flow splitting at an impacting tee, two-phase flows in tee junctions and

manifolds, slug flow and flow conditioning devices.

2.1 Splitting at Impacting Tee

Because of the geometry of the distribution manifold, two-phase flow coming

from the inlet wells impinges the opposite manifold wall, causing splitting of both liquid

and gas in an impacting configuration. Hong and Griston (1995) developed a model to

predict the liquid and gas splitting in impacting tees. The experimental facilities used in

their study included a splitting tee of the same diameter as the upstream pipe, and also

downstream separators to measure the liquid and gas that each branch carried.

For the majority of flow conditions tested by Hong and Griston, the flow pattern

upstream of the tee was stratified flow. Results from this experimental study cannot be

used in the analysis of the proposed distribution manifold. This is mainly due to the fact

that different flow patterns may exist in the different inlet wells connected to the

6

distribution manifold. Also, several impacting tees are found in a distribution manifold,

where the flow splitting in each tee may affect the flow behavior of others.

Wang and Shoji (2002) conducted an experimental study to evaluate the

fluctuating characteristics of two-phase flow splitting at a vertical impacting tee. This

study focused mainly on churn flow, which exhibits the strongest flow fluctuations as

compared to other flow patterns, such as bubble flow or annular flow. The main tube of

the impacting tee was vertical and the two branches were horizontal. Data included the

effect of the extraction flow ratios and the upstream superficial gas velocities on the flow

fluctuations.

2.2 Two-Phase Flow in T-Junctions

For single-phase flows, the present state of knowledge is sufficiently advanced to

enable the majority of cases of flow in junctions to be designed. In the case of two-phase

flow, however, the number of variables is much larger; in addition there are complicating

factors in the distribution and mixing of the phases. The problem is particularly acute for

dividing junctions, whereby either phase can pass preferentially into the side branch of

the junction. In any calculations involving junctions, the flows are governed not only by

what occurs at the junction itself but also by the flow resistances at different elements of

the entire system. Azzopardi (1992) showed that the liquid and gas flow split and the

pressure drop in a tee junction are affected by the angle between the main tube and the

side arm, the ratio of side arm to main tube diameters and the degree of rounding of the

corner. Because gravity can cause stratification, it is also necessary to specify the angle

7

between the main tube and the vertical, and the angle between the side arm and the main

tube.

The effect of the orientation of the side-arm in a tee junction, when the flow in the

main tube (horizontal) is stratified, was studied by Reimann and Khan (1984). They

found that initially only one phase is extracted, namely, the gas, when the side-arm is

vertical upwards, and liquid when side arm is vertical downwards. When the side-arm

was at the bottom of the main tube, Reimann and Khan observed that the free surface of

the liquid above the side-arm became depressed with some gas entrained into the side-

arm.

Taitel and Dukler (1987) analyzed the hydrodynamics near the exit of a pipe

carrying gas and liquid in horizontal stratified flow. This work demonstrated the effect of

pipe length on the stratified to non-stratified transition boundary. For low-viscosity

fluids, pipe length effects are unimportant but for high-viscosity liquids the transition

from stratified flow can be profoundly influenced by the pipe length.

2.3 Slug Flow

An important additional feature of the Distribution Manifold is its capacity to

minimize the effect of slug flow produced from wells, by providing damping time and

functioning as a pre-separator. Dukler and Hubbard (1975) developed the fundamental

and pioneering mechanistic model of slug flow. Based on their observations, Dukler and

Hubbard defined an idealized slug unit and suggested a mechanism for the flow. The

developed model is capable of predicting the slug hydrodynamic flow behavior, including

the length, holdup, velocity and pressure distributions.

8

Taitel and Barnea (1990) presented a comprehensive analysis of slug flow. It

aimed to extend the scope of slug flow modeling into a unified model for horizontal,

inclined and vertical upward flow. For the study of the hydrodynamics of the film,

several approaches were proposed, including rigorous, equilibrium film and open channel

flow approaches. Also, two methods to predict the pressure drop were described, namely,

a global force balance on a slug unit and a force balance on the liquid slug only. A

comprehensive discussion on closure relationship is also presented.

2.4 Manifolds

Manifold can be defined as more than one tee junction connected together to a

main pipe. Miller (1971), Collier (1976) and Coney (1980) conducted experiments on

manifolds. Miller (1971) noted that for single-phase flow, interaction between junctions

occurs if the interjunction distance is less than three main tube diameters. No equivalent

information was obtained for two-phase flow. However, some kind of interaction

between junctions for multiphase flow is expected, too. Collier (1976) analyzed a

horizontal system of four tubes linked by inlet and outlet headers. Void fractions and

pressure drop were used to determine the flow rates and qualities in each branch of the

manifold. It was found that the manifold results were consistent with the measurements

taken utilizing equivalent set of single junctions.

2.5 Flow Conditioning

Sarica et al. (1990) presented a mechanistic model for the prediction of the

required dimensions of a finger storage slug catcher. The approach is based on the effect

9

of the finger pipe diameter and inclination angle on the transition boundary between slug

flow and stratified flow. Based on this model, the required length and optimal downward

inclination angle of the fingers can be determined.

Ramirez (2000) presented experimental data on slug dissipation in helical pipes.

The data depict the effect of different operating conditions and helical pipe geometry on

slug dissipation. The data showed the effects of helix diameter, pitch angle and number of

circular turns on slug dissipation phenomenon.

Reinoso (2002) conducted a study on a novel device known as a slug damper. In

this work, experimental data were acquired in order to determine the slug propagation in

the slug damper and the corresponding outlet liquid flow rate, and to identify the

mechanisms involved in the slug damper performance. A mechanistic model was

developed for design purposes.

CHAPTER III

EXPERIMENTAL PROGRAM

3.1 Experimental Facility

An experimental flow loop has been constructed in the College of Engineering

and Natural Sciences Research Building, located in the North Campus of The University

of Tulsa. This indoor facility enables year around data acquisition and simultaneous

testing of different compact separation equipment. This chapter includes details of the

experimental facility, data acquisition procedure and representative experimental results.

Figure 3.1 shows a schematic of the experimental facility. The oil-water-air three-

phase flow facility is a fully instrumented state-of the-art, two-inch flow loop, enabling

testing of single separation equipment or combined separation systems. The three-phase

flow loop consists of a metering and storage section and a modular test section.

Following is a brief description of both sections.

Micromotion

Micromotion

DistributionManifold

AirFilter

Micromotion

3-PhaseSeparator

WaterTank

OilTank

GLCC/LLCC

LLHC

HPS

Micromotion

Micromotion

DistributionManifold

AirFilter

Micromotion

3-PhaseSeparator

WaterTank

OilTank

GLCC/LLCC

LLHC

HPS

Figure 3.1. Schematic of Experimental Facility

11

3.1.1 Metering and Storage Sections

Air is supplied from a compressor and is stored in a high-pressure gas tank. The

air flows through a one-inch metering section, consisting of Micromotion® mass flow

meter, pressure regulator and control valve. The liquid phases (water and oil) are pumped

from the respective storage tanks (400 gallons each), and are metered with two sets of

Micromotion® mass flow meters, pressure regulators and control valves. The pumping

station, shown in Figure 3.2, consists of a set of two pumps (10 HP and 25 HP equipped

with motor speed controllers) for each liquid-phase. Each set of pumps has an automatic

re-circulating system to avoid the occurrence of high pressure in the discharge line.

The liquid and gas phases can be either mixed at a tee junction and sent to the test

section or, in the case of the test section of the multiphase distribution manifold, the two

phases can flow separately through two 2” independent lines up to the test section.

Downstream of the test sections, the gas, oil-rich and water-rich streams flow through

three Micromotion® net oil computers to measure the outlet gas flow rate, and the flow

rate and water-cut of the two liquid streams. The three streams then flow into a three-

phase conventional horizontal separator (36-inch diameter and 10 feet long), where the

air is vented to the atmosphere and the separated oil and water flow back to their

respective storage tanks. A technical grade white mineral oil type, Tufflo® 6016, with a

specific gravity of 0.857 and a viscosity of 27 cp (@ 75 °F) is used as the experimental

oil along with tap water. For this research, only water and air were used in the

experimental program.

12

3.1.2 Modular Test Section

The metered three-phase mixture coming from the metering section can flow into

any of the six different test stations. This flexibility enables the testing of single

separation equipment, such as a GLCC, LLCC©2, Liquid-Liquid Hydrocyclone (LLHC),

Horizontal Pipe Separator (HPS©3), Multiphase Distribution Manifold or conventional

separators, or any combination of these, in parallel or series, forming a compact

separation system. Each test section is briefly described following.

GLCC/LLCC. This facility allows conducting experiments in a rudimentary

compact separation system using a Liquid-Liquid Cylindrical Cyclone (LLCC)

2 LLCC© Liquid-Liquid Cylindrical Cyclone-copyright, The University of Tulsa, 1998 3 HPS© Horizontal Pipe Separator-copyright, The University of Tulsa, 2001

Figure 3.2. Tanks, Pumping Station and Metering Section

13

downstream of a GLCC. These two separators working in series perform as a three-phase

separator.

Horizontal Pipe Separator (HPS). This facility allows conducting experiments

using a horizontal pipe separator to analyze the separation process between two

immiscible liquid phases (oil continuous).

Hydrocyclone (LLHC). This facility allows the conduction of experiments to

separate droplets of any liquid dispersed into another continuous liquid-phase. The main

condition to run this facility is that there must exist a difference in density between both,

continuous and dispersed phases.

Multiphase Distribution Manifold. This facility allows conducting experiments to

evaluate the gas and liquid split in a manifold for different flow configurations and

different flow conditions of the inlet wells. Downstream this manifold there are two

instrumented GLCCs enabling metering the gas and liquid phases flowing into each

separator, and subsequently, liquid and gas split ratios of the manifold can be determined.

3.1.3 Instrumentation, Control and Data Acquisition System

Control valves placed along the flow loop control the flow into the test sections.

The flow loop is also equipped with several temperature sensors and pressure transducers

for measurement of the in-situ temperature and pressure conditions. All output signals

from the sensors, transducers, and metering devices are collected at a central panel. A

state-of-the art-data acquisition system, built using LabView®, is used to both control the

flow into and out of the loop and to acquire data from the analog signals transmitted by

the instrumentation.

14

3.1.4 Multiphase Distribution Manifold/SlugDamper/GLCC System

The multiphase distribution manifold system, shown schematically in Fig. 3.3,

consists of four sections, namely, inlet wells, manifold, slug dampers, and down stream

GLCCs. Detailed description of each section is given below.

Inlet Well Section. The inlet wells are connected to the manifold from the liquid

and gas supply flow lines by two sets of four tees. Thus, four separate inlet gas lines and

four separate liquid lines are available, respectively. The flow rates in each of the 8 inlet

lines are measured using 8 rotameters. The inlet well section is shown in Figure 3.4. The

four inlet wells are simulated by combining pairs of single-phase liquid and single-phase

gas lines, resulting in 4 two-phase inlet wells, which are connected to the manifold.

Manifold Section. Figure 3.5 shows an isometric view of the distribution manifold

(dimensions are in inches). The manifold is an 8 feet long, 3- inch ID horizontal pipe

section. The four inlet wells are connected horizontally to the manifold with equal

spacing of 2 ft. The manifold has two 3” upper gas exits at the top and two 3” lower

liquid exits at the bottom, with a spacing of 4 ft. from each other. The upper and lower

exits each consist of a vertical section, 1 ft. long, which are radially opposed. The upper

exits are connected to the upper legs of the 2 slug dampers and the lower exits are

connected to the lower legs of the slug dampers. During operation, the production from

individual wells flows into the distribution manifold where a pre-separation of the phases

occurs, whereby the liquid goes to the lower exits and the gas goes to the upper exits.

15

Figure 3.3. Schematic of Manifold / Slug Damper / GLCC System

Gas Outlet

Vortex Meter

Vortex Meter

GLCC # 1 Rotameter

Gas Line

Liquid Line

Liquid Outlets to Micromotion

Slug Damper

GLCC # 2


WELL 4

WELL 3

WELL 2

WELL 1

16

When slugs or any increase in the liquid or gas flow rates from any well are

introduced into the manifold, they are dissipated and, therefore, the effect of these slugs

or perturbations on the system is minimized. It is important to note that because of the

geometry of the gas and liquid outlets and the way the manifold is designed, it must be

installed in an elevated position in such a way that upper and lower legs of the manifold

could be connected to the slug damper inlets. Doing so will allow the liquid and gas to

flow towards the GLCCs.

Figure 3.4. Liquid and Gas Inlet Flow Meters

17

Slug Damper Section. As shown in Figure 3.6, the slug damper consists of two

large diameter legs located one above the other. The two legs are 7.5 ft. long and 3- inch

diameter, where the lower leg is inclined downward at 1.4° while the upper leg is inclined

upward at 1.8°. The vertical distance between the two legs at the end of them is 28

inches. These two legs are connected to the GLCC, resembling a long dual GLCC inlet.

In this system there are two slug dampers connecting the distribution manifold with 2

downstream GLCCs. These 2 slug damper units allow the liquid and gas to run

independently from the manifold to the separators.

The main operational mechanism of the slug damper is a segmented orifice

located in the lower leg, just upstream of the GLCC. The orifice is open at the bottom and

Figure 3.5. Multiphase Distribution Manifold Schematic

(dimensions in inches)

18

closed at the top. When a slug hits the damper, due to the flow restriction provided by the

segmented orifice in the lower leg, the slug is damped, accumulating in the lower leg, and

fairly constant liquid flow rate enters the GLCC through the orifice. Note that in addition,

the distribution manifold provides an additional damping capacity upstream of the slug

dampers.

GLCC Section. The experimental facility includes 2 identical GLCCs, connected

to the 2 slug damper units. The GLCC, shown schematically in Figure 3.7, is a 6 feet

high, 3-inch ID vertical pipe, with dual inlets. The lower inlet of the GLCC is connected

to the lower leg of the slug damper. The GLCC inlet slot area is 25% of the inlet full bore

Figure 3.6. Slug Damper Manifold

19

cross sectional area and is connected tangentially to the vertical pipe. The upper inlet,

which is connected to the upper leg of the

slug damper, is a full bore pipe, connected

1 foot below the top of the vertical pipe.

The 2-inch ID GLCC gas outlet is located

radially at the top of the vertical GLCC

body. The 2-inch ID GLCC liquid outlet is

connected tangentially at the bottom of the

vertical GLCC section. Liquid and gas

flow rates from each GLCC are measured

separately and then recombined before

entering the 3-phase separator. Liquid

streams are measured using a

Micromotion® mass flow meter and gas

streams are measured using Vortex

shedding meter.

3.2 Experimental Setup

Over two hundred experimental test runs were carried out in this study to quantify

the performance of the distribution manifold, for different flow conditions and flow

patterns occurring in the inlet wells. These experimental tests include variations in gas

and liquid flow rates of each inlet well.

Figure 3.7. GLCC Separator Schematic

Liquid outlet

Gas outlet

Upper inlet (Gas)

Lower inlet (Liquid)

1’

1’

3.5’

0.5’

Liquid outlet

Gas outlet

Upper inlet (Gas)

Lower inlet (Liquid)

1’

1’

3.5’

0.5’

20

3.2.1 Parameter Definitions

Following is a description of the parameters used to determine the performance of

the multiphase distribution manifold.

Superficial Gas Velocity: The superficial gas velocity is defined as the in-situ

total volumetric gas flow rate in the manifold divided by the total cross sectional area of

the manifold. Combining the equation of state and the definition of gas superficial

velocity yields the value of this variable, as follows:

2**056.3

pG

GSG d

mVρ

= (3.1)

where,

SGV is the superficial gas velocity, in ft/s.

Gm is the gas mass flow rate, in lb/min.

Gρ is the density of the gas, in lb/ft3.

Pd is the pipe diameter, in inches.

Superficial Liquid Velocity: Similarly, the superficial liquid velocity is defined as

total liquid flow rate entering the manifold divided by the cross sectional area of the

manifold. Following is the equation used to determine the superficial liquid velocity.

2**056.3

pL

LSL d

mVρ

= (3.2)

where,

SLV is the superficial liquid velocity, in ft/s.

Lm is the liquid mass flow rate, in lb/min.

Lρ is the density of the liquid, in lb/ft3.

21

Pd is the pipe diameter, in inches.

Liquid Split Ratio: This is the ratio between the liquid flow rate at the exit of one

GLCC to the total liquid flow rate entering the manifold. This ratio is given in

percentage, as follows.

% 100*,

,

TOTALL

GLCCLL q

qSR = (3.3)

where,

LSR is the liquid split ratio, in %.

GLCCLq , is the liquid flow rate at the exit of any GLCC, ft3/min.

TOTALLq , is the total liquid flow rate entering the manifold, in ft3/min.

Gas Split Ratio: Similarly, the gas split ratio is the ratio between the gas flow rate

at the exit of one GLCC to the total gas flow rate entering the manifold, given in

percentage, as follows.

100*,

,

TOTALG

GLCCGG q

qSR = % (3.4)

where,

GSR is the gas split ratio, in %

GLCCGq , is the gas flow rate at the exit of any GLCC, in ft3/min.

TOTALGq , is the total gas flow rate entering the manifold, in ft3/min.

Gas Volume Fraction (GVF): The GVF is the volumetric fraction of gas flow rate

in the manifold. It is given by the ratio between the superficial gas velocity and the

mixture velocity inside the manifold.

22

Liquid Carry-Over (LCO): Operational condition for a given superficial gas

velocity and superficial liquid velocity, where some liquid is carried into the exit gas

stream. In the case of the manifold, liquid carry-over is defined when some liquid appears

in the horizontal section of either of the two upper gas legs of the manifold.

3.2.2 Test Procedure

The general procedure followed for conducting experiments is given below:

1. Start the liquid pump and the air compressor.

2. Fix the liquid and gas flow rates utilizing the flow control system, control

valve/Micromotion, using Labview program.

3. Measure the individual liquid and gas flow rates of the four inlet wells.

4. Measure the liquid and gas flow rates at the liquid and gas exits of the 2

down stream GLCCs.

5. Measure pressure and temperature of fluids in the system.

The detailed procedure and the purpose of each experiment are described in

section 3.3 of this chapter.

3.2.3 Test Matrix

The following data acquisition matrix was selected in order to study the

performance of the distribution manifold and the corresponding gas and liquid split

ratios.

23

Flow Configurations: Eight different flow configurations of the inlet wells were

analyzed. These flow configurations are shown schematically in Figure 3.8. In this figure,

L designates liquid dominant well and G designates that the well is gas dominant. The

eighth case, which is not shown in this figure, corresponds to equal inlet gas and liquid

flow conditions in all the wells.

Operating Flow Conditions: Air and water were used in this study. The ranges of

superficial velocities inside the manifold are:

• Superficial gas velocity: 10.5-30.5 ft/s.

• Superficial liquid velocity: 1.0-2.75 ft/s.

1 2 3 4

L GL L

CASE I

1 2 3 4

L LL G

CASE II

1 2 3 4

L GL G

CASE III

1 2 3 4

L L G G

CASE IV

1 2 3 4

G GL L

CASE V

1 2 3 4

L GG G

CASE VI

1 2 3 4

G GL G

CASE VII

Vsg: 10.5 fts/s to 30.5 ft/s Vsl: 1.0 ft/s to 2.75 ft/s

Figure 3.8. Flow Configurations in Inlet Wells

24

3.3 Experimental Results

In this section, detailed experimental results are presented. The results include:

• Operational envelope of Manifold/Slug Damper/GLCC integrated system

for liquid carry-over.

• Operational envelope of the Manifold itself for liquid carry-over.

• Liquid and gas split ratios in the Manifold.

• Liquid discharge resistance coefficient of the Manifold.

• Transient performance of the Manifold/Slug Damper/GLCC system under

liquid flow rate surges at the inlet.

When plotting the LCO envelope of either the manifold or the total system, the

superficial liquid velocity, Vsl, is plotted in the vertical axis and the superficial gas

velocity, Vsg, is plotted in the horizontal axis. When plotting the liquid or gas split ratios,

the gas volume fraction, GVF, in the distribution manifold is plotted in the horizontal

axis, and either the gas or liquid split ratio is plotted in the vertical axis.

3.3.1 Liquid Carry-Over Operational Envelope of Manifold / Slug Damper / GLCC

Integrated System

During normal operation, it is desirable to avoid liquid carry-over in either of the

two GLCCs, in order to maintain high separation efficiency. The operational envelope of

the Manifold/Slug Damper/GLCC System for Liquid Carry-Over is the onset of liquid

carry-over in either of the 2 downstream GLCCs. Operational envelopes can be presented

for either the GLCC or for the entire system to characterize their respective capacities to

ensure no liquid carry-over.

25

In the existing facility, it is possible to measure not only the total liquid and gas

flow rates entering the system, but also the gas and liquid flow rates into each GLCC.

Therefore the envelope of a single GLCC can be determined. Once the envelope for a

single GLCC is plotted, the operational envelope of the combined two identical GLCCs

operating in parallel can be plotted. So, the theoretical maximum capacity of the system

is twice the capacity of a single GLCC, if equal split from the manifold is provided.

Several experiments were conducted, with different inlet wells configurations, to

determine the operational envelopes for liquid carry-over of a single GLCC, two parallel

GLCCs and the envelope of the entire system (Manifold / Slug Damper / GLCCs), under

different flow conditions of the inlet wells. The procedure followed to determine these

operational envelopes is given below:

1. Choose the flow configuration of the inlet wells, as shown in Figure 3.8, and

adjust the flow through each well in such a way that in the adjustment process

the flow configuration is maintained, namely, either liquid or gas dominant or

equal flow.

2. Starting with high gas flow, fix the gas flow rate and increase the liquid flow

rate until liquid carry-over can be observed in either of the down stream

GLCCs.

3. Repeat step 2 with a lower gas flow rate.

Figure 3.9 shows the operational envelopes of the entire system under different

inlet well flow configurations. In this figure, the lower curve represents the envelope for a

single GLCC. The upper curve represents the operational envelope of two GLCCs

operating in parallel under equal flow conditions.

26

For flow configurations I, III, VI and VII, the wells that are liquid dominant or

gas dominant are placed in one end of the manifold, creating an uneven split. For these

cases, as shown in the Figure 3.9, for higher superficial gas velocities, the capacity of the

system is higher than the capacity of two parallel GLCCs. However, for lower superficial

gas velocities, the capacity of the system is reduced significantly. For cases I and VI, the

superficial gas velocity separating the 2 different capacity behaviors is around 26 ft/s. For

case III, this superficial gas velocity is around 23 ft/s and for case VII it is around 29 ft/s.

The explanation of the phenomenon is presented next.

Due to the uneven split generated at the Manifold, one GLCC can operate in one

edge of the individual GLCC envelope (low gas flow rate / high liquid flow rate) and the

Figure 3.9. Operational Envelope for LCO of Manifold / Slug Damper / GLCC System

Manifold / Slug DamperLiquid Carry-Over

0.00

0.50

1.00

1.50

2.00

2.50

3.00

3.50

0.00 5.00 10.00 15.00 20.00 25.00 30.00 35.00 40.00 45.00

Vsg (ft/s)

Vsl (

ft/s)

Single GLCC

Double GLCC

Case I

Case II

Case III

Case VI

Case VII

Case VIII Equal Flow

Single GLCC

2 Parallel GLCC's

Manifold/Slug Damper/GLCC's

27

other GLCC will operate on the opposite edge of the envelope (high gas flow rate / low

liquid flow rate). This uneven split can explain the fact that running the system under

uneven split and superficial gas velocities higher than 30 ft/s, the capacity of the system

can be higher as compared to the capacity of two GLCCs working in parallel.

For flow configuration II the envelope shows that as the superficial gas velocity

increases up to 30 ft/s, the capacity of the system approaches the theoretical capacity of

two parallel GLCCs operating under equal split conditions. Low superficial gas velocities

in this case greatly affect the capacity of the system, as can be seen in Figure 3.9.

Cases IV, V, which are not shown, and case VIII provide equal liquid and gas

flow rate in both GLCCs. Theoretically, the envelope for these cases should run parallel

to the envelope for two parallel GLCCs. The experimental data validate this hypothesis.

Due to the fact that these experiments were conducted under steady state

conditions and the level inside the GLCCs was almost constant and high enough from the

liquid outlet, no gas carry under was observed while conducting the experiments. It is

also important to mention that the capacity of the compressor limited the amount of

experimental data points that could be obtained. It was not possible to go beyond a total

superficial gas velocity of 30 ft/s.

3.3.2 Manifold Operational Envelope for Liquid Carry-Over

As defined before, the liquid carry-over for the manifold is defined as the locus of

all pairs of superficial liquid and superficial gas velocities that cause liquid carry-over in

the horizontal pipe of either of the two upper gas legs of the manifold. To obtain these

operational envelopes for LCO, the same procedure utilized to determine the envelope of

28

Figure 3.10. Operational Envelope of the Distribution Manifold

Manifold Operational Envelope for Liquid Carry-Over

0.0

0.5

1.0

1.5

2.0

2.5

3.0

5 10 15 20 25 30 35

Vsg (ft/s)

Vsl (

ft/s)

Case I

Case II

Case III

Case IV

Case V

Case VI

Case VII

Case VIII Equal Flow

the entire system was followed, and eight different envelopes were obtained for each of

the different inlet flow configuration, as shown in Figure 3.10.

It can be seen that even though there are differences in the envelopes for the

different configurations, the behavior of the manifold is fairly constant for superficial gas

velocities ranging from 18 to 25 ft/s, and superficial liquid velocities ranging from 1.5 to

2.0 ft/s.

By comparison of Figures 3.9 and 3.10, the operational envelopes for liquid carry-

over of the manifold for all the different flow configurations always fall below the

envelope of the entire system. For this reason, it is decided to use the envelope of the

manifold as a design criterion of the system to avoid liquid carry-over in the compact

separators. This will serve as a conservative approach for the design of the system. The

29

fact that the envelope of the manifold is chosen for the design and performance of the

entire system, and the fact that some liquid carry-over in the manifold does not

necessarily mean liquid carry-over in the downstream GLCCs, allows the designer to

establish a safety factor permitting slugs or sudden increment of gas or liquid flow rates

to enter the manifold without affecting the performance of the entire system.

3.3.3 Liquid and Gas Split Ratios in Distribution Manifold

The liquid and gas split ratios in the distribution manifold are two of the most

important parameters to be determined experimentally. To obtain the liquid and gas split

ratios in the manifold and how they are affected by variables such as the superficial gas

velocity, the superficial liquid velocity and inlet well flow configurations, the same eight

different cases, as presented in Figure 3.8, were analyzed. The procedure is described

below:

1. Choose a flow configuration for the inlet wells, according to Figure 3.8, and

adjust the rotameters in such a way that as total flow from the metering

section is changed in the adjustment process, flow through the individual

wells change in the same proportion, so that the desired flow configuration

remains the same.

2. Choose operational points that do not give liquid carry-over in the GLCCs,

starting from high gas flow rate and low liquid flow rate to low gas flow and

high liquid flow rate. For each operational point selected, individual liquid

and gas flow rates in both GLCCs were measured. As the total flow rate into

30

Figure 3.11. Liquid Split Ratio in the Distribution Manifold

Liquid Split ( G LCC# 1 over Total Flow ) vs. G VF

0.40

0.50

0.60

0.70

0.80

0.90

1.00

0.75 0.8 0.85 0.9 0.95 1

GVF

Liqu

id S

plit

Case I

Case II

Case III

Case IV

Case V

Case VI

Case VII

Case VIII Equal F low

the system is known, the liquid and gas split ratios can be calculated for each

separator.

The liquid and gas split ratios are plotted as a function of the gas volume fraction

(GVF) inside the manifold. Two different plots are obtained, which are described in the

following section.

Liquid Split Ratio: Figure 3.11 shows the liquid split ratio vs. the gas volume

fraction. It can be seen that in flow configurations I, III and VI, the liquid split ratio is

greatly affected by the GVF. For these cases, as GVF increases, the liquid split goes far

from 50%, which is the desirable value. In cases IV, V and VIII, theoretically the liquid

split ratio, as expected, is 50% because these flow configurations give equal gas and

liquid flow rate along the distribution manifold. Cases II and VII show a little deviation

with respect to 50% liquid split ratio value.

31

According to Figure 3.11 and the explanation given, cases I, III and VI should be

avoided. On the other hand, flow configurations similar to cases IV, V and VIII, which

give liquid split very close to 50 %, should be pursued. In cases where most of the wells

are liquid dominant or gas dominant, like cases II and VII, the liquid split ratio is

acceptable because it is close to 50%, regardless of the GVF.

Gas Split Ratio: Figure 3.12 shows how the measured gas split ratio is related to

the GVF. As can be seen, the gas split ratio in flow configurations I, III and VI is greatly

affected by the GVF. For these cases, as GVF decreases, the gas split ratio goes away

from 50%, which as mentioned before, is the desirable value. It also can be seen that all

other flow configurations, especially cases IV, V and VIII give a gas split ratio fairly

close to 50%, regardless of the GVF. Cases II and VII give a gas split ratio that has a

little deviation from the preferred 50% as GVF decreases.

Figure 3.12. Gas Split Ratio in the Distribution Manifold

Gas Split ( GLCC# 1 over Total Flow ) vs. GVF

0.00

0.10

0.20

0.30

0.40

0.50

0.60

0.75 0.8 0.85 0.9 0.95 1

GVF

Gas

Spl

it Case I

Case II

Case III

Case IV

Case V

Case VI

Case VII

Case VIII Equal F low

32

Comparing Figure 3.11 with Figure 3.12, it can be observed that cases I, III and

VI should be avoided while designing a distribution manifold, as for these cases the

liquid and gas split ratios are very sensitive to the GVF inside the manifold. In these

cases, low GVF values promote equal liquid split but also promote uneven gas split. Also

cases IV, V and VIII are the desirable flow configurations because the liquid and gas split

ratios are around 50%, regardless of GVF. Even though in cases II and VII liquid and gas

split ratios show a little deviation from 50%, these flow configurations are acceptable.

3.3.4 Resistance Coefficient of Manifold (Kl)

An important factor, which affects the liquid flow rate through the lower outlet

legs of the manifold, and in turn the capacity of the distribution manifold, is the

resistance of the outlet liquid leg. A high resistance to liquid flow will mean an increase

of the liquid level inside the manifold, resulting in lower capacity of the system. This is

also due to the fact that as the liquid level inside the manifold increases, the chance of

having liquid carry-over increases.

To enable prediction of liquid flow rate in each lower leg of the manifold, an

experiment was conducted to obtain an equation to relate the Resistance Coefficient (Kl)

of this liquid outlet to the Reynolds Number in the liquid legs. The procedure followed in

this experiment was to flow only liquid into the manifold, under even and uneven split,

measure the liquid level close to each liquid outlet, and obtain the corresponding liquid

flow rate through each leg from the GLCC outlet measurement.

33

Once these measurements were obtained, Bernoulli’s equation was applied to the

liquid flow in each leg to relate the potential energy available (expressed in terms of the

level inside the manifold) to the velocity of the liquid in the liquid leg. Bernoulli’s

equation applied to any stream line on the liquid surface level in the manifold gives:

.22

**2

12

1 VVKhg l =− (3.5)

where,

h is average liquid height in the manifold in the liquid outlet, in m.

1V is the velocity of the liquid in the liquid leg, in m/s.

lK is the resistance coefficient of the liquid leg.

Manipulating Equation 3.5 to obtain an equation for Kl, the following expression

is obtained,

1**22

1

−=V

hgKl (3.6)

Reynolds number in the liquid leg is calculated using the following equation:

µρ dVl **

Re 1= (3.7)

where,

lρ is the density of the liquid, in kg/m3.

1V is the velocity of the liquid in the liquid leg, in m/s.

d is the diameter of the liquid leg, in m.

µ is the viscosity of the liquid, in cp.

34

Once Resistance Coefficient and Reynolds number are calculated, a plot is

generated to relate these two variables, as given in Figure 3.13. This plot shows that as

Reynolds number increases, the Resistance Coefficient tends to approach a constant

value, as discussed next. This experiment was conducted in absence of gas flow, for this

reason, the effect of the gas or the effect of bubbles in this resistance coefficient was not

determined. Further analysis and experimentation would be required to determine the

resistance coefficient if the ratio of areas in the splitting tee is different than one and also

if the gas flow is considered important.

For low Reynolds numbers in the liquid leg, smaller than 30,000, the liquid flow

through the outlet is basically by free drainage, i.e., the connection manifold-liquid leg is

not totally covered with liquid. This phenomenon can be seen in sketch (a) of Figure

Figure 3.13. Resistance Coefficient for Liquid Outlet vs.

Reynolds Number

Kl vs. Reynolds

0

10

20

30

40

50

60

70

0 5,000 10,000 15,000 20,000 25,000 30,000 35,000 40,000 45,000

Reynolds Number

Kl

35

3.14. For Reynolds numbers bigger than 30,000, the liquid outlet of the distribution

manifold is totally flooded and the resistance coefficient tends to be constant and have a

value around 5. Sketch (b) of Figure 3.14 shows the level inside the manifold when the

inlet is totally flooded.

Results obtained in this study for the Resistance Coefficient are different from the

values found in the literature. Miller (1971) conducted some experiments for dividing

flow in 12 and 8 inch-ID tees, and reported a value around 2.0. Collier (1976) also

conducted experiments for standard tees and reported values around 1.3, regardless of the

diameter. Crane Corporation (1981) presents a Resistance Coefficient of 0.78 for a pipe

entrance that better represents a tee.

The difference between the values found in the literature and the experimental

results obtained are due to the fact that the geometry and the way the liquid flow towards

LIQ

WELLWELL

(b)

LIQ

WELLWELL

(a)

LIQ

WELLWELL

(b)

LIQ

WELLWELL

(a)

LIQ

WELLWELL

(a)

Figure 3.14. Conditions for Determination of Resistance Coefficient

36

the outlet is totally different for the different studies. Geometries studied by Miller (1971)

and Collier (1981) are tees, whereby the flow goes mainly through the run and some flow

goes through the side branch. In the case of the manifold, flow through the liquid leg

comes from both sides of the tee. Geometry analyzed by Crane Corporation represents a

pipe entrance from a big reservoir, where the flow could come from all points

surrounding the pipe entrance. In the case of the manifold, the only possible sources for

the liquid leg are the 2 sides of the tee.

3.3.5 Transient Performance of Multiphase Distribution Manifold

Several experiments were conducted to analyze the performance of the

distribution manifold under condition of transient flow and uneven liquid and gas split

ratios. Once this uneven split is reached, operate the system under the flow configuration

identified as case III in Figure 3.8. The procedure followed to conduct these runs is given

below:

1. Starting with no gas flow rate coming into the manifold and a given liquid

flow rate, wait until stable condition in the entire system, including the

manifold and in the GLCCs is reached.

2. Increase suddenly the total liquid flow rate entering the manifold and acquire

data of the liquid flow through each GLCC as a function of time.

3. Return to the original flow conditions.

4. Increase the gas flow rates, wait until stable condition is reached and repeat

steps 2 and 3.

37

No gas flow (Vsg = 0): In Figure 3.15, the LHS plot shows the performance of the

distribution manifold under transient conditions with no gas flow. The upper curve

represents the total liquid flow rate and the two lower curves represent the flow through

each downstream GLCC. The curve in the middle represents the total flow coming out of

the system. This plot shows that even though the inlet total flow is almost doubled, the

liquid flow in each GLCC changes very slowly, as compared with the step induced in the

total inlet liquid rate. The difference between the two upper curves, namely, the total inlet

flow rate and the total outlet flow rate represents the damping capacity of the system.

The RHS plot of Figure 3.15 shows the corresponding variation of the liquid split

ratio for GLCC # 2 for the transient conditions presented in the LHS of the figure. It can

be seen that for no gas flowing into the system, the liquid split ratio is not affected and is

essentially the same split obtained for stable flow conditions.

Figure 3.15. Transient Performance of Manifold Under No Gas Flow

Transient FlowLiquid Split in GLCC # 1 for Vsg=0 ft/s

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0 20 40 60 80 100 120 140

t (s)

Liqu

id S

plit

Transient PerformanceVsg=0 ft/s

0.00

0.50

1.00

1.50

2.00

2.50

0 20 40 60 80 100 120 140t (s)

Vsl (

ft/s)

Total Flow In

Total Flow Out

Flow in GLCC # 1

Flow in GLCC # 2

Transient FlowLiquid Split in GLCC # 1 for Vsg=0 ft/s

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0 20 40 60 80 100 120 140

t (s)

Liqu

id S

plit

Transient PerformanceVsg=0 ft/s

0.00

0.50

1.00

1.50

2.00

2.50

0 20 40 60 80 100 120 140t (s)

Vsl (

ft/s)

Total Flow In

Total Flow Out

Flow in GLCC # 1

Flow in GLCC # 2

38

Low gas flow (Vsg = 3.2 ft/s): Figure 3.16 shows the transient response of the

system under low gas flow into the manifold.

As can be seen, in the lower curves of the LHS plot of Figure 3.16, for low gas

flow rates entering the manifold, the flow through each GLCC increases almost as fast as

the increment seen in liquid legs for no gas condition.

Regarding the corresponding liquid split, it shows a constant and proper value for

most of the time during the duration of the experiment. When the liquid flow is increased

suddenly and when conditions return to the initial operational point, the liquid split ratio

shows a sudden increment, which could affect the performance of the system and the

performance of the metering system downstream the of GLCC’s.

Figure 3.16. Transient Performance of Manifold Under Low Gas Flow

Transient FlowVsg=3.2 ft/s

0

0.5

1

1.5

2

2.5

0 50 100 150 200t (s)

Vsl (

ft/s)

Total Flow In

Total Flow Out

Flow in GLCC # 1

Flow in GLCC # 2

Transient FlowLiquid Split in GLCC # 1 for Vsg=3.2 ft/s

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0 50 100 150 200

t (s)

Liqu

id S

plit

Transient FlowVsg=3.2 ft/s

0

0.5

1

1.5

2

2.5

0 50 100 150 200t (s)

Vsl (

ft/s)

Total Flow In

Total Flow Out

Flow in GLCC # 1

Flow in GLCC # 2


0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0 50 100 150 200

t (s)

Liqu

id S

plit

39

Moderate gas flow (Vsg = 6.7 ft/s): Figure 3.17 shows the transient response of

the system under a moderate gas flow rate in the manifold.

Figures 3.15, 3.16 and 3.17 show that as gas flow rate increases, the performance

of the distribution manifold and the liquid split under transient conditions are greatly

affected. For this reason and due to the absence of liquid level control valves for the

GLCCs in the experimental facility, it is difficult to maintain a constant liquid level in the

GLCCs and as a result only three experiments were conducted. It is recommended to

install liquid level control valves in the GLCCs for future testing.

Figure 3.17. Transient Performance of the Manifold Under Moderate Gas Flow

Transient PerformanceVsg=6.7 ft/s

0.00

0.50

1.00

1.50

2.00

2.50

0 50 100 150 200

t (s)

Vsl (

ft/s)

Total Flow In

Total Flow Out

Flow in GLCC # 1

Flow in GLCC # 2


0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0 50 100 150 200

t (s)

Liqu

id S

plit

Transient PerformanceVsg=6.7 ft/s

0.00

0.50

1.00

1.50

2.00

2.50

0 50 100 150 200

t (s)

Vsl (

ft/s)

Total Flow In

Total Flow Out

Flow in GLCC # 1

Flow in GLCC # 2


0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0 50 100 150 200

t (s)

Liqu

id S

plit

CHAPTER IV

MODEL DEVELOPMENT

A mechanistic model has been developed for predicting the hydrodynamic flow

behavior in the distribution manifold. This includes the liquid and gas split ratios in the

manifold for different flow conditions in each inlet well. This chapter includes the

“Design Model”, which is based on the Kelvin-Helmholtz stability analysis and the

“Manifold/Slug Damper/GLCC System Performance Model”, which is based on mass

balances, liquid level and pressure distributions inside the manifold, and flow patterns

and flow characteristics in each inlet well.

4.1 Manifold Design Model

The model developed for design purposes is divided into two sub-models:

• Manifold sizing sub-model for a given set of inlet wells configuration and

given liquid and gas flow rates.

• Liquid and gas outlets sizing sub-model.

4.1.1 Manifold Sizing

The manifold size is determined by ensuring stratification in the manifold. This is

carried out by applying a simplified Kelvin-Helmholtz stability analysis.

In general, the Kelvin-Helmholtz stability analysis deals with two fluid layers of

different densities ρ1 and ρ2 flowing horizontally with velocities v1 and v2, respectively.

The theory predicts whether a small disturbance on the surface will lead to the interface

being stable, with a wavy structure, or unstable with wave growth and, thus, destroying

41

the stratification between the two layers. The governing mechanisms, according to this

analysis, are: on the one hand, the gravity and surface tension forces tend to stabilize the

flow; on the other hand, the relative motion between the two layers creates a suction

pressure force, owing to the Bernoulli effect, tending to destroy the stratified structure of

the flow. Figure 4.1 shows the case of a stationary finite wave on the gas-liquid interface

in the manifold. For the simplified stability analysis the surface tension effect is

neglected.

The stabilizing gravity force acting on the wave is,

( )( ) θρρ Cosghh GLGG −− ' (4.1)

where,

Gh is the height of the gas phase inside the pipe, in m.

'Gh is the distance from top of the pipe to the top of the bubble, in m.

θ is the angle between the pipe and the horizontal plane, in degrees.

The pressure suction force causing wave growth is given by,

( )22''

21

GGG vvPP −=− ρ (4.2)

Figure 4.1. Stationary Finite Wave on Gas-Liquid Interface

42

where,

P is the pressure inside the pipe, in Pa.

'P is the pressure where the bubble occurs, in Pa.

Gv is the velocity of the gas phase inside the pipe, in m/s.

'Gv is the velocity of the gas phase where the bubble occurs, in m/s.

From the continuity relationship,

''GGGG AvAv = (4.3)

where,

GA is the area of the gas phase inside the pipe, in m2.

'GA is the area of the gas phase where the bubble occurs, in m/s.

The transition from stratified to non-stratified flow regime takes place when the

suction force is greater than the gravity force. Combining Equations 4.1 through 4.3, the

criterion for unstable stratified flow structure is,

21

)(

−>

IG

GGLG S

ACosgcvρ

θρρ (4.4)

where,

+

=

G

G

G

G

AA

AA

c'

2'

2

12 (4.5)

43

For low liquid level in the pipe GG AA ≈' , therefore c=1. Similarly, for high liquid

level in the pipe 0' →GA , and hence c=0. On the basis of these boundary conditions it is

hypothesized that,

dhc L−= 1 (4.6)

where,

Lh is the height of the liquid phase inside the pipe, in m.

Substituting Equation 4.6 into Equation 4.4 results in the final criterion for this

transition boundary,

( ) 21

1

−

−>

IG

GGLLG S

ACosgdhv

ρθρρ (4.7)

where,

IS is the length of the interface gas-liquid, in m.

If the gas velocity inside the manifold (LHS) is greater than the expression on the

RHS, then the Bernoulli suction force overcomes the gravity force causing the flow to be

unstable. To ensure stratification in the manifold, a sufficiently large manifold diameter is

chosen, which can guarantee that the RHS of Equation 4.4 will be bigger than the given

gas velocity inside the manifold, resulting in stable stratified configuration.

This stability analysis is applied to determine the optimum manifold diameter

using two criteria, as given below.

Criterion 1: Figure 4.2 shows a schematic of the distribution manifold and the

procedure used to determine its diameter. As shown in the figure, the manifold is divided

44

into several sections. In this criterion each section has only one well connected and the

diameter of each section of the manifold is determined independently using the liquid and

gas flow rates of the well connected, and applying Equation 4.7 to ensure stratification.

The effect of the wells in neighboring sections is neglected.

Criterion 2: As shown schematically in Figure 4.3, with this criterion, interaction

between the flow in the different sections is considered. Thus, the diameter of a section is

determined taking into consideration the effect of the wells from neighboring sections. It

is assumed that flow rates of wells located in inner sections (see figure) split equally

between the section on the left and the section on the right. Thus, for example, the

calculation of the diameter of section 2 is based on the flow rate of the well connected to

this section, plus half of the flow rates of the wells connected to the inner section.

Section 1 Section 2 Section 3 Section 4

Section 1 Section 2 Section 3 Section 4

Figure 4.2. Manifold Sizing-Criterion 1

45

The manifold diameter is chosen as the larger diameter calculated using both

criteria. Liquid and gas flow rates through each outlet leg are calculated based on the

assumption that wells on inner sections of the manifold flow towards the closest outlet. In

case that the number of wells in the inner section is odd, liquid and gas rates from this

well will split equally to the outlets. Figure 4.4 shows how split is calculated. Thus, the

diameter of the distribution manifold, as well as the liquid level and pressure distribution

along the manifold can be determined.

Section 1 Section 2 Section 3 Section 4Inner Section

Section 1 Section 2 Section 3 Section 4Inner Section

Figure 4.3. Manifold Sizing-Criterion 2

Figure 4.4. Liquid and Gas Split Calculations

46

4.1.2 Gas and Liquid Outlets Sizing

Once the liquid level and pressure distributions inside the manifold are calculated

and liquid and gas flow rates through each liquid and gas legs are known, liquid and gas

outlets diameters can be determined. The model for the prediction of the liquid and gas

outlet diameters is given next.

Liquid Leg Diameter: This model assumes that the potential energy possessed by

any particle on the free surface of the liquid, is transformed into velocity and head losses

due to the liquid leg entrance. This can be expressed in the terms of resistance coefficient.

Figure 4.5 shows the variables involved in this model.

Applying Bernoulli equation to the particle located on the free liquid surface

above the liquid outlet,

gvHf

gvh

22

222

1

21 =−+ (4.8)

where,

LIQ

WELLWELL

h

dL

V1

V2

LIQ

WELLWELL

h

dL

V1

V2

Figure 4.5. Variables Involved in Liquid Outlet Sizing

47

h is the height of the free surface level, assuming stratified flow, in m.

1v is the velocity of the particle in the free surface level, in m/s.

2v is the velocity of the liquid in the liquid leg, in m/s.

2

1Hf is head loss in the pipe entrance, in m.

The head loss is given by the following equation,

gvKHf L 2

222

1= (4.9)

where LK is the liquid outlet resistance coefficient, determined experimentally, whose

values is assumed to be constant, namely, LK = 5.

Assuming that the velocity of the particle in the free surface of the liquid, 1v , is

negligible as compared with the velocity of the particle in the outlet leg 2v , and

combining Equations 4.8 and 4.9 yields,

LKhgv

+=

12

2 (4.10)

With the outlet liquid flow rate, and the liquid velocity calculated using Equation

4.10, the area of the liquid outlet is calculated as:

2vQA L

L = (4.11)

where,

LQ is the liquid flow rate in the lower liquid outlet leg, in m3/s.

LA is area of the lower liquid outlet leg, in m2.

Once the area of the lower liquid leg outlet is calculated, the diameter can be

obtained in a straightforward manner as:

48

πL

LA

d4

= (4.12)

where Ld is the diameter of the lower liquid outlet leg, in m.

The same procedure can be applied to the other liquid outlets.

Gas Leg Diameter: To calculate the upper gas leg outlet diameter, the model

assumes that the pressure drop along the gas leg is equivalent to one velocity head. Thus,

to determine the velocity of the gas in the upper leg the following equation is used:

1

2

2 =∆

GGG

vK

Pρ

(4.13)

where,

P∆ is the pressure drop, assumed to be equivalent to 1 velocity head, in Pa.

Gv is the velocity of the gas in the upper gas outlet leg, in m/s.

Gρ is the density of the gas, in kg/m3.

GK is the gas outlet resistance coefficient, obtained from the literature (Crane CO,

1981).

Manipulating Equation 4.13, the velocity of the gas in each gas leg, can be

obtained, as follows,

GGG K

Pvρ∆

=2 (4.14)

With the gas flow rate and the velocity calculated using Equation 4.14, the area of

the upper gas outlet leg is determined as:

49

G

GG vQ

A = (4.15)

where,

GQ is the gas flow rate in the upper gas outlet leg, in m3/s.

GA is area of the pipe in the upper gas leg, in m2.

Once the area of the upper gas outlet leg is calculated, its diameter can be

obtained, namely,

πG

GA

d4

= (4.16)

where,

Gd is the diameter of the upper gas outlet leg, in m.

The same procedure can be repeated for the other gas outlet.

4.2 Manifold / Slug Damper / GLCC System Performance Model

The model developed for performance prediction is based on the Hardy-Cross

Method (Streeter, Wylie and Bedford, 1997) for both liquid and gas phases. The Hardy-

Cross method uses the liquid height or flow rate as boundary conditions, carrying energy

balances on the system, whereby either the liquid height or flow rate can be calculated.

This method can be easily applied to closed systems, like aqueducts, or open systems like

the distribution manifold. Hardy- Cross method usually neglects the kinetic energy of the

fluid because the predominant factor is the potential and pressure energy

+ ρ

PZ . In

this study, however, the kinetic energy of the fluids is not neglected because it is of the

same order of magnitude as the potential and energy pressure term.

50

Even though friction losses in the system are very small, those are considered in

present study because if fluids involved have high viscosity and distance between wells

in the manifold are long, then frictional losses become significant.

Figure 4.6 presents a schematic of the manifold, which shows all the variables

involved in the proposed mechanistic model.

The variables involved are as follows:

Ql 1,2,3,4 is the liquid flow rate in each well, in m3/sec.

Qg 1,2,3,4 is the gas flow rate in each well, in actual m3/sec.

Ql a, Ql d , are the liquid flow rates in each liquid outlet leg, in m3/sec.

Figure 4.6. Variables Involved in Manifold Performance Model

Qg1 Qg2 Qg3 Qg4

Ql1 Ql2 Ql3 Ql4

Qla

a d

Qga

Qld

Qgd

b c

Pknown

Qg1 Qg2 Qg3 Qg4

Ql1 Ql2 Ql3 Ql4

Qla

a d

Qga

Qld

Qgd

b c

Pknown

51

Qg a, Qg d , are the gas flow in each gas outlet leg, in m3/sec.

Ql t, is the total liquid flow rate, in m3/sec.

Qgt , is the total gas flow rate, in m3/sec.

a, b, c , d, are the nodes considered in the model.

Pknown , is the known pressure where both gas streams combine, in Pa.

Assumptions made to develop the mechanistic model are as follow:

• Friction factor is calculated assuming smooth pipe.

• Pressure drop for the gas phase in the GLCCs, due to the swirling motion,

is assumed constant and expressed in terms of equivalent length of pipe.

• Pressure drop in the slug damper, for both, liquid and gas phases, is

calculated assuming single phase flow.

The proposed model needs an initial guess of the liquid and gas flow rates in

either node “a” or node “d”. Next, applying Hardy-Cross method, the liquid heights in the

different nodes, the flow rates through the different nodes, and the liquid and gas flow

rates through the remaining outlets are calculated. Since liquid height along the manifold

changes, the area available for the gas change, so pressure distribution changes, too. This

pressure distribution works as the driving force for the liquid. Figure 4.7 shows the

simplified piping and nodes scheme of the manifold, indicating the liquid and gas flows

inside the manifold.

52

Once initial guess for liquid and gas flow rates through the liquid and gas outlets

in node “a”, (Qla and Qga) is made, mass balance is applied for both phases, liquid and

gas, inside the manifold, following the equations described below:

1llalab QQQ −= (4.17)

1ggagab QQQ −= (4.18)

labllbc QQQ −= 2 (4.19)

gabggbc QQQ −= 2 (4.20)

lbcllcd QQQ += 3 (4.21)

gbcg QQQ += 3gcd (4.22)

lcdlld QQQ += 4 (4.23)

a b c d

QlaQld

Ql1

Qg1

Ql2Qg2

Ql3Qg3

Ql4

Qg4

QgaQgd

Qgab

Qlab

Qgbc

Qlbc

Qgcd

Qlcda b c d

QlaQld

Ql1

Qg1

Ql2Qg2

Ql3Qg3

Ql4

Qg4

QgaQgd

Qgab

Qlab

Qgbc

Qlbc

Qgcd

Qlcd

Figure 4.7. Simplified Piping and Nodes Model

53

gcd4 QQQ ggd += (4.24)

4321 llllldlalt QQQQQQQ +++=+= (4.25)

4321 gggggdgagt QQQQQQQ +++=+= (4.26)

Since the model attempts to solve the liquid and gas phases simultaneously,

determination of the liquid level in the distribution manifold has to be carried out first.

This first calculation is done solving the model considering only liquid inside the

manifold. The procedure involved in this calculation is given below.

4.2.1 Calculation of liquid height at node “a”

This height is calculated using the equation for resistance coefficient determined

experimentally, as plotted in Figure 3.13. First step is to calculate the velocity in the

liquid outlet in node “a”.

lo

lala A

Qv = (4.27)

( )lalaa K

gv

h += 12

2

(4.28)

Reynolds number in the liquid outlet can be calculated using Equation 3.7, then;

3741.16 Re*10*6277.9 −=lK (4.29)

where,

lav is the velocity in the liquid outlet leg at node “a”, in m/s.

loA is area of the pipe in the liquid outlet, in m2.

laK is the Resistance Coefficient in the liquid outlet in node “a”, dimensionless.

54

ah is the liquid height in the node “a”, in m.

4.2.2 Calculation of Head Losses between nodes

The head losses between nodes are calculated using the Darcy-Weisbach equation

(Streeter, Wylie and Bedford, 1997). In the case of nodes “a” and “b” this equation is:

gvv

dL

fHf abab

m

tb

a 2= (4.30)

where,

b

aHf is the liquid or gas head losses between nodes “a” and “b”, in m.

f is the Moody friction factor, dimensionless.

abv is the velocity of either liquid or gas phase between nodes “a” and “b”, in m/s.

md is the diameter of the manifold, in m.

tL is the distance between nodes, in m.

The Moody friction factor is calculated assuming smooth pipe, using the Blasius

equation (Mounson, Young and Okiishi, 2002), as follows:

nCf −= Re (4.31)

where C = 64 and n= 1 for laminar flow and C = 0.184 and n= 0.2 for turbulent flow.

Reynolds number is a function of the physical properties of the fluid, the liquid or

gas hydraulic diameter and the velocity of the fluid in either the liquid or gas cross

sectional area. The hydraulic diameter for liquid (Sl) and gas (Sg) and liquid (Al) and gas

(Ag) cross sectional areas are calculated following the Taitel and Dukler model (1976),

which assumes stratified flow in the pipe. The present model considers not only the

55

magnitude of the velocity but also the direction of the fluid. Equation 4.27 takes into

consideration this fact.

Once head losses between nodes are calculated, energy balances are applied

between the nodes, as follows:

( ) ( )

gv

gP

hHfg

vg

Ph ba hlab

l

gbb

b

alhlab

l

gaa 22

22

++=+++ρρ

(4.32)

( ) ( )

gv

gP

hHfg

vg

Ph bc hlbc

l

gbb

c

blhlbc

l

gcc 22

22

++=+++ρρ

(4.33)

( ) ( )

gv

gP

hHfg

vg

Ph cd hlcd

l

gcc

d

clhlcd

l

gdd 22

22

++=+++ρρ

(4.34)

where, ),,,( dcbagP are the pressures in the different nodes, in Pa.

Pressure at the gas discharge downstream of the GLCCs and the physical

configuration of the system are known. Since the gas mass balance has already been

maintained, the pressure at the beginning of the upper gas legs downstream of the

manifold can be easily calculated, using the following equations:

2

2_

__ 2 go

ga

o

aeqaggknownega A

QdL

fPP

+= ρ (4.35)

2

2_

__ 2 go

gd

o

deqdggknownegd A

QdL

fPP

+= ρ (4.36)

where,

egaP _ and egdP _ are the pressures of the gas at the beginning of the upper legs,

downstream of the manifold, in Pa.

gρ is the density of the gas, in kg/m3.

56

agf _ and dgf _ are the friction factors in legs a and d, respectively, dimensionless.

aeqL _ and deqL _ are the equivalent lengths of pipe and accessories in the upper

gas legs, in m.

goA is the gas outlet area, in m2.

gaQ and gdQ are the gas flow rates through nodes “a” and “d”, respectively, in

m3/s.

Pressure in the internal nodes of the manifold is calculated by applying energy

balance on the gas-phase, as follows,

( )( )

( ) 2

2

_2

21

21

2 go

gagagega

hg

gabgabggga A

QKP

AQQQ

Pa

++=+

+ ρρ (4.37)

( )( )

( ) 2

2

_2gcdgcd

24

21

2 go

gdgdgegd

hg

gggd A

QKP

AQQQ

Pd

++=+

+ ρρ (4.38)

( ) ( )ba hg

gabgabggb

b

aghg

gabgabgga A

QQPHf

AQQ

P 22 22ρρ

+=++ (4.39)

( ) ( )cd hg

ggc

d

cghg

ggd A

QQPHf

AQQ

P 2gcdgcd

2gcdgcd

22ρρ

+=++ (4.40)

Applying this mass and energy balances throughout all the nodes, liquid level

distribution along the manifold is obtained. Height in node “d” must be compared with

the one obtained applying Equations 4.27, 4.28 and 4.29 for the liquid-phase in node “d”.

Iterative procedure, changing either liquid or gas flow rates through nodes “a” or “d”,

57

must be done until the difference between these two values satisfies a convergence

criterion. A detailed description of the procedure is explained below:

1. Based on well arrangement, assume a value for liquid and gas split ratio for

either node “a” or “d”.

2. Applying Equations 4.17 to 4.26, determine the liquid and gas flow rates in

all nodes.

3. Calculate the liquid level inside the manifold, assuming no gas flow. This

calculation is used as a first approximation to determine the cross sectional

area available for the gas phase in all nodes.

4. Calculate gas pressure at the inlet of the upper gas legs of the slug damper,

downstream of the manifold, applying Equations 4.35 and 4.36.

5. Using Equations 4.37, 4.38, 4.39 and 4.40, calculate the pressure in the

internal nodes. Head losses for liquid and gas phases must be calculated

using Equations 4.30 and 4.31.

6. Once pressure distribution along the manifold has been determined,

calculate the liquid height in all nodes. Equation 4.28 is used for nodes “a”

and “d” and Equations 4.32, 4.33 and 4.34 must be used to calculate liquid

heights in nodes “b”, “c” and “d”, respectively. Since liquid height in node

“d” is calculated via energy and mass balances, these two values must be

equal. If these two values do not satisfy a convergence criterion, then a new

value for liquid and gas split must be assumed and the entire procedure from

step 1 is repeated.

58

Above discussion summarizes the model development in this study. Comparison

of the model prediction and experimental data is given in the next chapter.

CHAPTER V

SIMULATIONS AND RESULTS

This chapter presents a comparison between the model predictions and the

experimental data for the outlet (downstream of the GLCCs) liquid and gas split ratios.

The comparison is carried out for different flow configurations and different liquid and

gas flow rates at the inlet wells. The liquid and gas split ratios comparison is presented as

a function of the GVF. In all figures shown in this chapter, liquid and gas outlet split

ratios are related to the GLCC No. 2, as identified in Figure 3.3. Please refer to Figure 3.8

for identifying the different inlet well configurations.

5.1. Liquid Split Ratio

The developed mechanistic model is tested against the acquired experimental data

for outlet liquid split ratio (related to GLCC No. 2 in Figure 3.3), for several inlet flow

configurations. Figure 5.1 presents a comparison between the model predictions versus

experimental data for outlet liquid split ratio, for inlet wells flow configuration I. In this

experiment, total superficial liquid velocity ranges from 1.1 to 2.8 ft/s, while the

superficial gas velocity ranges from 10 to 30 ft/s. In this flow configuration, three out of

the four wells are liquid dominant and only one, the well located in one edge of the

manifold, is gas dominant. This configuration was maintained throughout the

experimental runs, regardless of the variation in the total liquid and gas entering the

distribution manifold. Good agreement between the model and the experimental data is

observed for this flow configuration. It can also be seen that the deviation between the

60

model predictions and the experimental data increases as GVF increases. However, this

deviation is around ±6% for high GVF, which is still low and can be considered as a good

agreement. For low GVF this deviation is much smaller.

Model predictions and experimental data are compared in Figure 5.2 for liquid

split ratio, in inlet wells flow configuration III for which the highest uneven liquid split

ratio occurs. In this flow configuration, two wells on one end of the distribution manifold

are liquid dominant and the other two are gas dominant. Good agreement between model

predictions and experimental data is observed for this case. As can be seen, the deviation


Liquid Split Ratio in Flow Configuration I

Inlet Flow Configuration I

0.50

0.55

0.60

0.65

0.70

0.75

0.85 0.87 0.89 0.91 0.93 0.95 0.97 0.99

GVF

Liqu

id S

plit

Rat

io

Experimental

Model

+ 5%

- 5%

1 2 3 4L GL L1 2 3 4L GL L


0.50

0.55

0.60

0.65

0.70

0.75

0.85 0.87 0.89 0.91 0.93 0.95 0.97 0.99

GVF

Liqu

id S

plit

Rat

io

Experimental

Model

+ 5%

- 5%

1 2 3 4L GL L1 2 3 4L GL L

61

between experimental data and model prediction is lower than ±8% for all points

analyzed. This deviation remains fairly constant regardless of the GVF.

Figure 5.3 shows the comparison between the model prediction and the

experimental data for inlet wells configuration VI. In this flow configuration, three of

four wells are gas dominant and only one, the well located in one end of the manifold, is

liquid dominant. Good agreement between model predictions and experimental data is

observed for this flow configuration. It can be seen that deviation is smaller than 5% for

all flow conditions. As can be seen, for superficial gas and liquid velocities, which give


Liquid Split Ratio in Flow Configuration III

Inlet Flow Configuration III

0.50

0.55

0.60

0.65

0.70

0.75

0.80

0.84 0.86 0.88 0.9 0.92 0.94 0.96 0.98 1

GVF

Liqu

id S

plit

Rat

io

Experimental

Model

+ 8%

- 8%

1 2 3 4L GL L1 2 3 4L GL G


0.50

0.55

0.60

0.65

0.70

0.75

0.80

0.84 0.86 0.88 0.9 0.92 0.94 0.96 0.98 1

GVF

Liqu

id S

plit

Rat

io

Experimental

Model

+ 8%

- 8%

1 2 3 4L GL L1 2 3 4L GL G

62

GVF higher than 0.91, the model under predicts the liquid split ratio, while for GVF

smaller than 0.91 the model over predicts the data.

The comparison between the model predictions and the experimental results for

inlet wells flow configuration IV is given in Figure 5.4. In this well configuration, the

two wells located in both ends of the distribution manifold are liquid dominant and the

two wells located in the inner section are gas dominant. This particular configuration

promotes equal liquid and gas split ratio, as was described in Chapter 3. Good agreement

Figure 5.3. Comparison between Model Predictions and Experimental Data for Liquid

Split Ratio in Flow Configuration VI

Inlet Flow Configuration VI

0.50

0.55

0.60

0.65

0.70

0.75

0.84 0.86 0.88 0.9 0.92 0.94 0.96 0.98 1

GVF

Liqu

id S

plit

Rat

io

Experimental

Model

+ 5%

- 5%

1 2 3 4L GL L1 2 3 4L GG G


0.50

0.55

0.60

0.65

0.70

0.75

0.84 0.86 0.88 0.9 0.92 0.94 0.96 0.98 1

GVF

Liqu

id S

plit

Rat

io

Experimental

Model

+ 5%

- 5%

1 2 3 4L GL L1 2 3 4L GG G

63

is observed between model predictions and experimental data for this inlet wells flow

configuration, too. Note that the deviation obtained for values of GVF around 0.88 and

larger than 0.95 seems to be due to experimental data error. Nevertheless, deviation

between experimental values and model prediction is smaller than ±10% for all points

evaluated for this case.


Liquid Split Ratio in Flow Configuration IV

Inlet Flow Configuration IV

0.40

0.42

0.44

0.46

0.48

0.50

0.52

0.54

0.56

0.58

0.60

0.8 0.82 0.84 0.86 0.88 0.9 0.92 0.94 0.96 0.98 1

GVF

Liqu

id S

plit

Rat

io

Experimental

Model

+ 5%

- 5%

1 2 3 4L GL L1 2 3 4L LG G


0.40

0.42

0.44

0.46

0.48

0.50

0.52

0.54

0.56

0.58

0.60

0.8 0.82 0.84 0.86 0.88 0.9 0.92 0.94 0.96 0.98 1

GVF

Liqu

id S

plit

Rat

io

Experimental

Model

+ 5%

- 5%

1 2 3 4L GL L1 2 3 4L LG G

64

5.2 Gas Split Ratio

The developed mechanistic model was tested against the acquired experimental

data for outlet gas split ratio for the same inlet wells flow configurations used for liquid

split ratio evaluation. Figure 5.5 presents a comparison between the model predictions

and experimental data for outlet gas split ratio for inlet wells flow configuration I. Model

predictions show a good agreement with experimental results for GVF higher than 0.92

(less than 5% error). However, for low superficial gas velocities, namely, low GVF,

deviation between model prediction and experimental results increases to about ±10%.


Gas Split Ratio in Flow Configuration I


0.25

0.30

0.35

0.40

0.45

0.50

0.55

0.85 0.87 0.89 0.91 0.93 0.95 0.97 0.99

GVF

Gas

Spl

it R

atio

Experimental

Model + 5%

- 5%

1 2 3 4L GL L1 2 3 4L GL L


0.25

0.30

0.35

0.40

0.45

0.50

0.55

0.85 0.87 0.89 0.91 0.93 0.95 0.97 0.99

GVF

Gas

Spl

it R

atio

Experimental

Model + 5%

- 5%

1 2 3 4L GL L1 2 3 4L GL L

65

Figure 5.6 shows the comparison between model predictions and experimental

results for gas split ratio in inlet well flow configuration III. As can be seen, model

predictions for gas split ratio in this configuration, show a very good agreement with the

data, with deviation no larger than ±3% for all points evaluated.

The comparison for gas split ratio for inlet well flow configuration VI is shown in

Figure 5.7. Model predictions show good agreement for GVF higher than 0.93. For

values lower than this, predictions show a considerable deviation. As GVF decreases, this


Gas Split Ratio in Flow Configuration III


0.25

0.30

0.35

0.40

0.45

0.50

0.84 0.86 0.88 0.9 0.92 0.94 0.96 0.98 1

GVF

Gas

Spl

it R

atio

Experimental

Model

+ 3%

- 3%

1 2 3 4L GL L1 2 3 4L GL G


0.25

0.30

0.35

0.40

0.45

0.50

0.84 0.86 0.88 0.9 0.92 0.94 0.96 0.98 1

GVF

Gas

Spl

it R

atio

Experimental

Model

+ 3%

- 3%

1 2 3 4L GL L1 2 3 4L GL G

66

deviation seems to increase, it can be seen in Figure 5.7, but still seems to be an

acceptable agreement within ±20%.

For inlet wells flow configuration IV, where equal liquid and gas split is expected,

model predictions show a very good agreement with experimental data, as presented in.

Figure 5.8. The agreement for this case is within ±5%, except for GVF ranging from 0.9

to 0.94.


Gas Split Ratio in Flow Configuration VI


0.25

0.30

0.35

0.40

0.45

0.50

0.55

0.60

0.84 0.86 0.88 0.9 0.92 0.94 0.96 0.98 1GVF

Gas

Spl

it R

atio

Experimental

Model

- 15%

+ 15%

1 2 3 4L GL L1 2 3 4L GG G


0.25

0.30

0.35

0.40

0.45

0.50

0.55

0.60

0.84 0.86 0.88 0.9 0.92 0.94 0.96 0.98 1GVF

Gas

Spl

it R

atio

Experimental

Model

- 15%

+ 15%

1 2 3 4L GL L1 2 3 4L GG G

67


Gas Split Ratio in Flow Configuration IV


0.40

0.42

0.44

0.46

0.48

0.50

0.52

0.54

0.56

0.58

0.60

0.8 0.82 0.84 0.86 0.88 0.9 0.92 0.94 0.96 0.98 1

GVF

Gas

Spl

it R

atio

Experimental

Model

+ 5%

- 5%

1 2 3 4L GL L1 2 3 4L LG G


0.40

0.42

0.44

0.46

0.48

0.50

0.52

0.54

0.56

0.58

0.60

0.8 0.82 0.84 0.86 0.88 0.9 0.92 0.94 0.96 0.98 1

GVF

Gas

Spl

it R

atio

Experimental

Model

+ 5%

- 5%

1 2 3 4L GL L1 2 3 4L LG G

CHAPTER VI

CONCLUSIONS AND RECOMMENDATIONS

This study investigated theoretically and experimentally a novel flow

conditioning device, namely, the Multiphase Distribution Manifold, to be used upstream

of compact separation systems. The following are accomplished during this study:

• A 3” ID Multiphase Distribution Manifold facility has been designed,

constructed, instrumented and installed in the indoor loop located in North

Campus of The University of Tulsa. Two slug dampers and two GLCCs are

attached downstream of the Distribution Manifold.

• Over 200 experimental runs were conducted. Eight different inlet well

configurations were evaluated. The superficial liquid and gas velocities

inside the 3” ID manifold were: VSL from 1.1 to 2.8 ft/s and VSG from 10 to

30 ft/s.

• For each experimental run, the measured data included: total liquid flow

rate, total gas flow rate, liquid and gas flow rate through each inlet well,

liquid and gas flow rate through each GLCC, static pressure in the system,

and temperature.

• Data analysis shows that the Distribution Manifold, for different flow

configurations used, is successful in providing a good distribution of liquid

and gas flow rates for the downstream separators.

69

• Data analysis shows that there are several inlet well configurations that

promote a better distribution of liquid and gas flow rate, more than other

configurations.

• Inlet well configurations that promote uneven split or maldistribution of the

liquid and gas flow rates can provide liquid and gas capacities in the system,

higher than the capacity achieved with two parallel GLCCs working under

equal split conditions.

• The operational envelope of the Distribution Manifold could be used as a

design criterion of the system. This serves as a conservative approach for the

design of the entire system, allowing establishing a safety factor permitting

slugs or sudden increments of gas or liquid flowing rates into the manifold

while avoiding upset of the downstream GLCC separators.

• Some experiments were conducted to test the capacity of the Distribution

Manifold under transient flow, increasing suddenly the total liquid entering

the Distribution Manifold. Preliminary results show that the manifold, in

conjunction with the slug damper, is successful in dissipating transient

slugging or sudden increments in the liquid flow rates, and provides fairly

constant flow rate into the GLCCs.

• A mechanistic model was developed for the prediction of the hydrodynamic

flow behavior in the Distribution Manifold. This model is based on the

Hardy-Cross method, which uses the liquid height or flow rate as boundary

conditions, carrying energy balances on the system. This is a non-linear and

70

iterative model. The model enables the prediction of the liquid and gas split

ratios in the manifold, as well as in the downstream GLCC separators.

• Comparison between the model predictions and the acquired experimental

data reveals a good accuracy for liquid and gas split ratios in all flow

configurations tested. Deviation between prediction and experimental data

generally ranges from 3% to 15%.

The following are recommended for future studies:

• Install level control valves in the liquid legs of the GLCCs to allow

conducting experiments in transient flow covering a wider range of liquid

and gas flow rates. This will allow the enhancement of dampening capacity

of the Distribution Manifold.

• The mechanistic model was implemented using the software Mathcad,

whereby the procedure to obtain the liquid and gas split ratios is an iterative

procedure. It is recommended to write the code using a programming

language, such as FORTRAN or VISUAL BASIC, which could give a faster

convergence.

• The Distribution Manifold is intended to work in conjunction with the Slug

Damper. It is recommended to combine the model presented by Reinoso

(2002) for the slug damper with the model presented in this study to predict

the performance of the entire system.

• Develop a design procedure and specific design criteria for field

applications of Distribution Manifold and integrate this into TUSTP GLCC

design software.

71

• Develop a mechanistic model for slug flow stability analysis and validate

this model using experimental data.

NOMENCLATURE

Symbols

A Cross sectional area, L2, m2

C Blassius coefficient

d Diameter, L, m

f Moody friction factor

F Head losses, L, m

g Acceleration due to gravity, (9.81) L/t2, m/s2

GVF Gas Volume Fraction

h Height in the manifold, L, m

Hf Head losses in pipe entrance, L, m

K Resistance Coefficient

L Length, L, m

LCO Liquid Carry-Over

m Mass Flow Rate, m/t, lbm/min

n Blassius coefficient

p Pressure, m/l t2, Pa

Q, q Flow Rate, L3/t, m3/s

Re Reynolds Number

S Hydraulic diameter, L, m

SR Split Ratio

t, T Time, s

73

v Velocity, L/t, m/s

z Potential energy, m L2/t3, W

Greek Letters

∆ Differential

ρ Density, m/L3, lbm/ft3

µ Viscosity, m/Lt, cp

π 3.14159265

θ Angle of inclination of pipe, °

Subscripts

eq equivalent

G, g Gas

L, l Liquid

m Manifold

o Outlet

P Pipe

SG Superficial Gas

SL Superficial Liquid

REFERENCES

1. Azzopardi, B.: “Two-Phase Flows at T-Junctions,” International Journal of

Multiphase Flow, v.18, n.6, Nov. 1992, pp. 677-713.

2. Collier, J.G.: “Single-Phase and Two-Phase Flow Behavior in Primary Circuit

Components,” Proceedings of NATO Advanced Study Institute, Istambul,

Turkey, 1976.

3. Coney, M. W.: “Two-Phase Flow Distribution in a Manifold System,” European

Two-Phase Flow Group Meeting, Glasgow, 1980.

4. Crane CO.: “Flow of Fluids Through Valves, Fittings and Pipe,” Technical Paper

No. 410, 1981.

5. Dukler, A. and Hubbard, M.: “A Model for Gas-Liquid Slug Flow in Horizontal

and Near Horizontal Tubes,” Ind. Eng. Chem. Fund., 14, 1975, pp. 337-347.

6. Hong, K.C. and Griston, S.: “Two-Phase Flow Splitting at an Impacting Tee,”

SPE Production & Facilities, v.10, n.3, Aug. 1995, pp. 184-190.

7. Miller, D.S.: “Internal Flow, A Guide to Losses in Pipe and Duet Systems,”

BHRA Cranfield, Bedford, UK, 1971.

75

8. Munson, B., Young, D. and Okiishi, T.: “Fundamentals of Fluid Mechanics,”

2002, 4th Edition, John Wiley & Sons.

9. Ramirez, R.: “Slug Dissipation in Helical Pipes,” M.S. Thesis, The University of

Tulsa, 2000.

10. Reimann, J.and Khan, M.: “Flow through a Small Break at the Bottom of a Large

Pipe with Stratified Flow,” 2nd Int. Topical Meeting on Nuclear Reactors

Thermohydraulics, Santa Barbara, California, 1984.

11. Reinoso, A.: “Design and Performance of Slug Damper,” M.S. Thesis, The

University of Tulsa, 2002.

12. Sarica, C., Shoham, O. and Brill, J. P.: “A New Approach for Finger Storage Slug

Catcher Design,” 1990 Offshore Technology Conference (OTC), Houston, Texas,

May 7-10.

13. Streeter, V., Wylie, B. and Bedford, K.: “Fluid Mechanics,” 9th Edition 1997,

McGraw Hill.

14. Taitel, Y. and Barnea, D.: “Two-Phase Slug Flow,” Academic Press Inc. 1990.

76

15. Taitel, Y. and Dukler, A.: “Effect of Pipe Length on the Transition Boundaries for

High-Viscosity Liquids,” International Journal of Multiphase Flow, v.13, n.4, Jul-

Aug 1987, pp. 577-581

16. Wang, S. and Shoji, M.: “Fluctuation Characteristics of Two-phase Flow Splitting

at a Vertical Impacting T-junction,” International Journal of Multiphase Flow,

v.28, n.12, December, 2002, pp. 2007-2016.

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