Gas Lift

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Study of the dynamics, optimization and selection of intermittent

gas-lift methods—a comprehensive model

Odair G. Santos a , Sergio N. Bordalo b,*, Francisco J.S. Alhanati c

a PETROBRAS, Brazil b Department of Petroleum Engineering, State University of Campinas, SP Caixa Postal 6052, 13083-970 Campinas, SP, Brazil

cCFER, Brazil

Abstract

The gas-lift is a widely used method of artificial lift, for which there are different design options. There are empirical rules of

thumb to choose between the continuous gas-lift (CGL) and the intermittent gas-lift (IGL), but little exists in the literature for

the selection among the different intermittent gas-lift designs. Furthermore, computer simulators for these processes usually are

not in the public domain. This work presents a numerical model to study the behavior of the Conventional IGL, the IGL with

chamber (IGLC), the IGL with plunger (IGLP) and the IGL with pig. Simulations are presented under various reservoir

conditions, for different settings of the operation’s parameters. The model’s results can aid the engineer in the determination of

the optimum values of the parameters for each design option and in the choice of the most adequate IGL design for a particular

well. D 2001 Published by Elsevier Science B.V.

Keywords: Gas-lift; Petroleum production; Computer simulation; Optimization

1. Introduction

Artificial lift is used in petroleum production when

the energy of the reservoir is not enough to sustain the

flow of oil in the well up to the surface with satisfac-

tory economic return. Selection of the proper artificial-

lift method is critical to the long-term profitability of the oil well; a poor choice will lead to low production

and high operating costs. There is very little margin for

error when one is designing lift systems for petroleum

fields. However, proper selection of the best method

still is based on past experience, strong opinions, fa-

miliarity of operating personnel with the equipment,

preferences of company experts that favor some meth-

od, and unsubstantiated technical myths. There is a

strong need for reliable procedures of selection and

design. Computer programs that simulate the operation

of lift systems are an important part of such procedures

and the main concern of the present work.

The gas-lift is a widely used method of artificial lift,where gas is injected in the production well providing

energy to the flow. In some instances, the continuous

injection of gas, named continuous gas-lift (CGL), is

not efficient, and the intermittent (periodical) injection

of gas, named intermittent gas-lift (IGL), becomes the

more economical alternative. Different design options

are available to implement the IGL: conventional IGL,

IGL with plunger (IGLP), IGL with chamber (IGLC)

and IGL with pig, also known as pig-lift (PL). There

are some empirical, though questionable, rules of

0920-4105/01/$ - see front matter D 2001 Published by Elsevier Science B.V.

P I I : S 0 9 2 0 - 4 1 0 5 ( 0 1 ) 0 0 1 6 4 - 4

* Corresponding author. Fax: +55-19-2894916.

E-mail address: [email protected] (S.N. Bordalo).

www.elsevier.com/locate/jpetscieng

Journal of Petroleum Science and Engineering 32 (2001) 231–248

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thumb to choose between the CGL and IGL (Table 1),

but little exists in the literature for the selection among

the different IGL designs. Furthermore, there are no

public models or comparative studies of the dynamicsof all the IGL process cycles that are consistent for all

designs (there are computer codes that are property of

private or government companies, and that were

developed for one design option only). The engineer

needs a tool to determine the performance of the

various designs under certain field conditions, to tune

the operational parameters to their optimum values,

and to choose the best IGL method. The purpose of

this paper is to provide such a tool to rationalize these

tasks.

This paper introduces a computer model to simulate

the conventional IGL, IGLP, IGLC and PL. All the

IGL methods work in cycles, and each cycle is made of

stages the follow transient flow processes, conse-

quently, the task of simulation is non-trivial. Besides,

vertical two-phase flow of oil and gas occurs in the

well, which adds to the complexity of the computa-

tions. In the IGLP system, a solid plunger separates the

oil and gas flowing in the well, to prevent fallback of

oil. In the IGLC system, a chamber is used to accu-

mulate the oil at the bottom of the well, reducing the

back-pressure against the reservoir formation. In the

PL system, a foam-pig separates the oil and gas flow-ing in the well to prevent fallback of oil, and a double

column is used inside the well.

The computer simulation employs sets of time-

dependent differential equations that govrn the vari-

ous phases of each IGL cycle. Although each IGL

design has its own idiosyncrasy, they also share some

common structure and processes which are modeled

and dealt with in a way that is consistent across all

designs.

Examples are presented for specific conditions of

the petroleum reservoir, and for optimization of the

operational parameters. The performance of the differ-

ent IGL designs are compared on the basis of an eco-

omic criterion — with each design operating at its

optimum point.

1.1. Literature review

Clegg et al. (1993) presented an extensive over-

view of artificial lift design considerations, compari-

son of methods and their normal operating conditions.

Chacın (1994) discussed the state of the art of the de-

sign of IGL methods, presented a simplified algorithm

for the calculation of the production rate, and a proce-

dure to select the best IGL method—according to his

criteria, the one with the greater ratio of produced oil

volume to gas injected volume. Brown and Jessen

(1962), Brill et al. (1967), and Neely et al. (1974) did

some experimental work on specific field installations

of conventional IGL, establishing empirical rules for

the setting of the operational parameters. Although

they provide useful guidelines, those rules lack in gen-

erality. White (1963) developed the first simple math-

ematical relationships for the conventional IGL and

did experiments on laboratory installations. Machado

(1988) developed a mechanistic model coupling phys-

ical principles and empirical correlations to calculate

some variables of the IGL system. Liao (1991) ob-tained theoretical results that showed good agreement

with Brown, Brill and Neely. White (1982) conducted

tests with and without a plunger to demonstrate the

reduction of liquid fallback in the plunger case. Mower

et al. (1985) used different plungers to study the effect

of plunger geometry on the fallback. Chacın et al.

(1992) developed a mechanistic model, introducing

the empirical findings of Mower and Lea into the IGLP

model. Brown described the advantages of IGL with

chamber for reservoirs with low static head and low

productivity index, and provided a simplified proce-dure to estimate the average flow pressure at the bot-

tom of the well. Winkler and Camp (1956) applied the

IGL with chamber to reservoirs of low static head but

high productivity index. Berdeja and Mariaco (1971)

discussed the principles that should be applied to the

analysis of the IGL with chamber, optimization of the

method, and listed results of field cases. Acevedo and

Cordero (1991) presented field experiments showing

an increase of oil production and decrease of gas con-

sumption. The pig-lift was developed in one petroleum

Table 1

Practical criteria for selection of continuous (CGL) and intermittent

(IGL) gas-lift

Static head (hs/ H w) Productivity index (PI, m3/day MPa)

High

( ! 20)

Medium

(5 to 20)

Low

(V5)

High ( ! 0.7) CGL CGL/IGL IGL

Medium (0.4 to 0.7) CGL/IGL CGL/IGL IGL

Low (V0.4) IGL IGL IGL

O.G. Santos et al. / Journal of Petroleum Science and Engineering 32 (2001) 231–248232

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company somewhat recently, consequently, very few

publications exist. Lima and Cardoso (1993) devel-

oped a simulator to study the potentiality of this new

technique. Lima (1996) presented the working principles of the pig-lift, and argued in favor of its advan-

tages over the other IGL methods.

Although the present model is inspired on Liao’s

work on the conventional IGL, it includes several orig-

inal elements that make it quite different and complete,

such as the liquid film and the fallback; moreover, it

has the generality necessary to deal with all IGL de-

signs.

2. IGL systems

2.1. The conventional IGL

Fig. 1a illustrates the conventional IGL cycle and

its stages. The motor valve controls the injection of gas

following a timer program. The gas-lift valve is set to

open at a certain pressure in the casing; at this point the

tubing is already loaded with a column a liquid. The

gas elevates the liquid slug, leaving behind a film of

liquid— the fallback. At some point, the motor valve is

closed. The slug is produced at the surface, and after

that, the gas is produced also; some liquid is produced

by the dragging of the liquid film and droplets dis- persed in the gas. The gas-lift valve closes and the

decompression of the tubing begins. Finally, the res-

ervoir pressure feeds the bottom of the well, reloading

the system for the next cycle.

The main parameters of the system are shown in

Table 2, with their values for the simulation run of the

case-example for this paper. The following operating

parameters were tuned to find the optimum economic

daily rate of produced oil volume to injected gas vol-

ume: V gi — injected gas volume per cycle; l si — initial

length of the liquid slug; P co —casing pressure to open

the gas-lift valve. Besides, the values of the reservoir

static pressure P r and reservoir productivity index PI

were varied to study the suitability of the IGL.

2.2. The IGL with plunger

Fig. 1b illustrates the IGLP cycle and its stages. The

scheme is similar to the IGL, except that, in this case,

the gas pushes the plunger up elevating the liquid slug.

Fig. 1. (a) The conventional intermittent gas-lift cycle. (b) The intermittent gas-lift with plunger cycle. (c) The intermittent gas-lift with chamber

cycle. (d) The intermittent gas-lift with pig cycle.

O.G. Santos et al. / Journal of Petroleum Science and Engineering 32 (2001) 231–248 233

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The fallback of liquid is admitted to be null for an ideal

plunger (no leakage). During the reloading of the well

for the next cycle, the plunger plummets back to its

initial position.

For the sake of comparison with the IGL, the same

well/reservoir system described in Table 2 was consi-

dered. Also, the same operating parameters, as in the

case of the IGL, were adjusted for the optimization of

the cycle: V gi, I si, P co.

2.3. The IGL with chamber

Fig. 1c illustrates the IGLC cycle and its stages. The

scheme is similar to the IGL, except that, when the gas-

lift valve opens, first the gas pushes the liquid from the

chamber into the tubing, then it elevates the liquid

slug. Fallback is also observed in the IGLC. During

reloading, the reservoir fluid is accumulated in thechamber, while in the IGL, the fluid fills the tubing.

Again, the well/reservoir system of Table 2 was

used to define the work conditions. The operating pa-

rameters adjusted for optimization are V gi and l si.

2.4. The pig-lift

Fig. 1d illustrates the PL cycle and its stages. As

with the other cycles, a motor valve controls the in-

jection of gas following a timer program, but, in this

case, a double-tubing is loaded with columns of liquid.The gas pushes the pig, transferring the liquid to one

leg, and then elevating the liquid slug. The fallback of

liquid is admitted to be null for an ideal pig. The slug is

produced at the surface, and after that, the gas is pro-

duced also. The motor valve is closed, and decom-

pression of the tubing begins. Finally, the reservoir

pressure feeds the bottom of the well, reloading both

legs of the tubing for the next cycle while the pig sinks

to its ready position. The functions of the tubing legs

are exchanged at every cycle.

Once again, the simulations were run for the well/

reservoir system of Table 2. The only operating para-

meter adjusted for optimization of the pig-lift is l si.

3. IGL model

3.1. System of equations

Table 3 shows the basic equations used in the simu-

lation of all the IGL methods. The equations are de-

rived from fundamental mass and momentum balances

applied to subsystems of the IGL (more details are

found in Santos, 1997, including the closure equa-

tions—equations of state, flow rate equations, valve’s

equation, geometric relationships, two-phase flow cor-

relations, head-loss correlations). Basic-equation b1

simulates the pressurization of the casing, B-equation b3 simulates the gas mass inside the tubing, B-equa-

tion b5 simulates the liquid film left behind by the slug,

B-equation b7 simulates the fallback, B-equations b8

and b9 simulate the slug elevation and B-equation b13,

in each of its forms, simulates the loading of the slug

for each IGL method.B-equation b3.1 simulates the

transfer stage of the IGLC, while B-equations b16.1

and b16.2 simulate the transfer stage of the pig-lift. For

each stage of the cycle, a system of non-linear time-

differential equations (Eq. (1)) was formed following

prescription in Table 4. The systems were solved withthe aid of commercial software of mathematical tools.

Xmvar

i¼1

ani

d X i

dt ¼ bn; n ¼ 1::meq : ð1Þ

3.2. Simulation

Fig. 2 shows examples of the simulation output for

the case-example described in Table 2, with a reser-

voir pressure P r = 8.35 MPa and a productivity index

Table 2

System parameters for the simulation of the case example

H w well depth 1500 m P wh well-head pressure 0.7 MPa co oil sp. gravitya 0.825

Dc casing diameter 5 1/2 in. P gi compressor pressure 7.0 MPa cw water sp. gravitya 1.07

Dt tubing diameter 2 3/8 in. T s surface temperature 27 °C dg gas rel. density b 0.7

H gv depth of gl-valve 1480 m aT temperature gradient 0.031 K/m uw water volume fraction 50%

a At standard conditions, relative to standard water. b At standard conditions, relative to standard air.


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Table 3

Basic equations for the IGL, IGLC, IGLP and PL

Variables—A: code variable; A; flow area; D: code variable; F ; flow switch; H: code variable; P: code variable; P ; pressure; Q: code variable; T ;

temperature; V : volume; Z : real gas factor; e: head-loss; f : friction factor; g : gravity; h: height; q: standard flow rate; t : time; v : velocity; w:

weight; y: film thickness; q: density.

Indexes—P: plunger/pig; Lx: leg-x; c: casing; ch: chamber; b: bottom; f: film; g: gas; h: head; i: injection valve; l: liquid; p: produced; r:

reservoir; s: surface; sc: standard condition; t: tubing; v: gas-lift valvep; w: well; c1: casing at inj-valve/injection leg top; c2: casing at gl-valve/

injection leg pig; c3: chamber at gl-valve; c4: chamber at gas–liq contact; t1: tub glV/orif./U-bott; t2: tub gas–slug contact; t3: tubing at slug-

top; t4: tub liquid column top; ts: tubing at surface.

V cdqgc

dt ¼ q

gscð F giqgi À F gvqgvÞ ð b1Þ

P c2 ¼ P c1 expð H gv=kÞ ð b2Þ

Achhch

dqgch

dt þ

d

dt ðqgt AgHgÞ

¼ qgscðQ À F gpqgpÞ ðb3Þ

Ach

d

dt

Àqgchðhch À hgÞ

Á¼ qgsc F gvqgv ð b3:1Þ

PA À PB ¼ f gt qgt

v 2gt

2

Hg

Dt

þ qgt g Hg ð b4Þ

P c3 À PD ¼ qgch g ðHgchÞ ð b4:1Þ

P c1 À P c2 À f gt qgt

v 2gt

2

Hg

Dt

þ qgt g Hg ¼ 0 ð b4:2Þ

P c1 À P gt1 À f gtqgL1

v 2gt

2

H t

Dt

þ q gL1 gH t ¼ 0 ð b4:3Þ

P gt1 À P t2 À f gtqgL2

v 2gt

2

hg

Dt

À qgL2 ghg ¼ 0 ð b4:4Þ

Hf

d Af

dt ¼ F rqr À qf ð b5Þ

P wb ¼ P t1 þ ql g ð H w À HoÞ ð b6Þ

Atv l À Agv gt ¼ Af v f ð b7Þ

dhl

dt ¼ fhl < Ho : v l; 0g ð b8Þ

dhg

dt ¼ fhg < Ho : v gt; 0g ð b9Þ

dhg

dt ¼ fhg > 0 : Àv gt; 0g ð b9:1Þ

ðHl À hgÞdv l

dt þ v 2l À 1 À

Ag

At

v 2f À

Ag

At

v 2gt

À1

ql

P t2 À PC Àwp

At

þ g ðHl À hgÞ

þ f lv 2l

2

Hl À hg

Dt

þ ewh ¼ 0 ðb10Þ

PC ¼ P wh þ qgts g ðHo À HlÞ ð b11Þ

Af

dv f

dt þ 2p

Dt

2À y

v f

d y

dt À

f g

8

qgt

ql

v 2gt

Hg

H t1

þf l

4v 2f p

Dt

2À

Af

H t1

P t1 À P wh

ql

þ gAf ¼ 0 ðb12Þ

d

dt ð Ach þ AgÞhlr ¼ ~ F rqr þ F f qf ð b13Þ

P t1 À P E ¼ ql ghlr ð b14Þ

v l At À v gt AG ¼ qr ð b15Þ

hg

dv gt

dt À

ðPF À P t1Þ

ql

þ f lv 2gt

2

hg

DÀ ghg ¼ 0 ð b16:1Þ

hl

dv l

dt þ v l

AG

At

v lAG

At

À 1

þ 2

qr

At

À

ð P t1 À P t3Þ

ql

þ f lv 2l2

hl

Dt

þ ghl ¼ 0 ðb16:2Þ


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Table 4

(a) Description of equation set for each stage of the IGL cycle

Stage Initial event Equation set

Injection injection valve opens: F gi = 1 [b1,b2] F gi = 1, F gv = 0.

Elevation gas-lift valve opens: F gv = 1 [b1,b2,b3,b4,b5,b6,b7,b8,b9,b10,b11]

F gi = 1, F gv = 1, F gp = 0, F r =1, PA= P t1, PB = P t2, PC = P t3, Hg = hg, Hf = hg, Hl = hl, qgt =qgt1, v gt = v b, qf = qfs.

Production slug reaches the surface: h1 = H gv [b1,b2,b3,b4,b5,b6,b7,b9,b10]

F gi = 1, F gv = 1, F gp = 0, F r =1, PA= P t1, PB = P t2, PC = P wh, Hg = hg, Hf = hg, Hl = H gv, q¯gt =qgt1, v gt = v b, qf = qfs.

Decompression I gas reaches the surface: hg = H gv [b1,b2,b3,b4,b5,b6,b12]

F gi={1;0}, F gv = 1, F gp = 1, F r =1, PA= P t1, PB = P wh, Hg = H gv, Hf = H gv, qgt = qgt1s, v gt = 0.5v gs, qf = qfs.

Decompression II gas-lift valve closes: F gv = 0 [b3,b4,b5,b6,b12,b13,b14]

F gv = 0, F gp = 1, F r = 0, F ˜ r = 1, F f =0, PA= P t4, PB = P wh, Hg = H gv À hlr , Hf = H gv À hlr , qgt = qgt4s, v gt =0.5v gs, qf = qfs.

Decompression III film flow reverses: v f = 0 [b3,b4,b5,b6,b13,b14]

F gv = 0, F gp = 1, F r = 0, F ˜ r = 1, F f =1, PA= P t4, PB = P wh, Hg = H gv À hlr , Hf = H gv, qgt = qgt4s, v gt =0.5v gs, qf = qfb.

Loading gas pressure relieved: v gs = 0 [b5,b6,b11,b13,b14]

F r

= 0, F ˜ r

= 1, F f

=1, PC

= P t4

, Hl

= hlr

, Hf

= H gv

, qf

= qfb

.

Cycle restarts liquid reloaded: hlr = l si *loop back*

(b) Description of equation set for each stage of the IGLP cycle


Injection injection valve opens: F gi = 1 [b1,b2]

F gi = 1, F gv = 0.

Elevation gas-lift valve opens: F gv = 1 [b1,b2,b3,b4,b6,b8,b9,b10,b11,b13,b14]

F gi = 1, F gv = 1, F gp =0, PA= P gt1, Pb = Pt2, Pc = P t3, Hg = hg, Hl = hl, qgt =qgt1, v gt = v l, F ˜ r = 1, F f = 0.

Production slug reaches the surface: hl = H gv [b1,b2,b3,b4,b6,b9,b10,b13,b14]

F gi = 1, F gv = 1, F gp =0, PA= P gtl, PB = P l2, Pc = P wh, Hg = hg, Hl = H gv, qgt = qgt1, v gt = v l, F ˜ r = 1, F f = 0.

Decompression I gas reaches the surface: hg = H gv [b1,b2,b3,b4,b6,b13,b14]

F gi={1;0}, F gv = 1, F gp =1, PA= P gtl, PB = P wh, Hg = H gv, qgt = qgt1s, v gt =0.5v gs, F ˜ r = 1, F f = 0.

Decompression II gas-lift valve closes: F gv = 0 [b3,b4,b6,b13,b14] F gv = 0, F gp = 1, P A= P t4, PB = P wh, Hg = H gvÀ hlr , qgt = qgt4s, v gt =0.5v gs, F ˜ r = 1, F f = 0.

Loading gas pressure relieved: v gs = 0 [b6,b11,b13,b14]

PA= P t4, PC = P t4, Hl = hlr , F ˜ r = 1, F f = 0.

Cycle restarts liguid reloaded: hlr = l si *loop back*

(c) Description of equation set for each stage of the IGLC cycle


Injection injection valve opens: F gi = 1 [b1,b2]

F gi = 1, F gv = 0.

Transfer gas-lift valve opens: F gv =1 [b1,b2,b3.1,b4.1,b6,b8,b9.1,b11,b15,b16.1,b16.2]

F gi = 1, F gv =1, Pc = P t3, PD = Pc4, Hl = h1, Hgch = hch À hg, qgch = qgc34, v gt = v gch.

Elevation chamber unloaded: hg = 0 [b1,b2,b3,b4,b4.1,b5,b6,b7,b8,b9,b10,b11]

F gi = 1, F gv = 1, F gp = 0, F r = 1, PA= P t1, PB = P t2, PC = P t3, PD = P t1, Hg = hg, Hf = hg, Hl = hl, Hgch = hch, qgch = qgc35, qgt = qgt1,V gt = V b, qr = qfs.Production slug reaches the surface: ht = Htl [b1,b2,b3,b4,b4.1,b5,b6,b7,b9,b10]

F gi = 1, F gv = 1, F gp = 0, F r = 1, PA= P tl, PB = P t2, Pc = P wh, PD = P tl, Hg = hg, Hf = hg, Hl = H tl, qr = qfs, Hgch = hch, qgch = qgc35, qgt =qgt1, v gt = v b.

Decompression I gas reaches the surface: hg = H tl [b1,b2,b3,b4,b4.1,b5,b6,b12]

F gi={1;0}, F gv = 1, F gp = 1, F r =1, PA= P tl, PB = P wh, PD = P tl, Hg = H tl, Hf = H tl, Hgch = hch, qgch = qgc35, qgt = qgt1s, v gt =0.5v gs, qr = qfs.

Decompression II gas-lift valve closes: F gv = 0 [b3,b4,b4.1,b5,b6,b12,b13,b14]

F gv = 0, F gp = 1, F r = 1, F ˜ r = 1, F f =0, PA= Pt4, PB = P wh, Hg = H tl À hlr , Hf = H tl À hlr , qgch =qgc3, qgt = qgt4s, v gt =0.5v gs, qr = qfs.

Decompression III film flow reverse V f = 0 [b3,b4,b4.1,b5,b6,b13,b14]

F gv = 0, F gp = 1, F r = 0, F ˜ r = 1, F f =1, PA= Pt4, PB = P wh, Hg = H tl À hlr , Hf = H tl, qgch =qgc3, qgt = qgt4s, v gt = 0.5v gs, qf = qfb.

Loading gas pressure relieved: V gs = 0 [b5,b6,b11,b13,b14]

F r = 0, F ˜ r = 1, F f =1, PC = P t4, Hl = hlr , Hf = Htl, qf = qfb.

Cycle restarts Liquid reloaded: hlr = l si At /( At + Ach) * loop back *


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P I =1 0 m3/day MPa; this pressure corresponds to a

ratio of static head to well depth hs/ H w = 0.6. The gas-lift valve has a seat diameter d gv = 0.5 in. The opration

parameters of the IGL and IGLP tuned for optimiza-

tion, as mentioned in Section 2, are represented here

by the non-dimensional trio {l si/ hs, P to/ P co,V gi/ V gref },

defined as:

l si/ hs, ratio of the initial length of the liquid slug

to the reservoir static head; P to/ P co, ratio of tubing pressure to casing pres-

sure to open the gas-lift valve;

V gi/ V gref , ratio of the gas volume injected per cycle to the reference volume.

The reference volume is the required volume calcu-

lated by Liao’s formula. For the cases of the conven-

tional IGL and the IGL with plunger, shown,

respectively, in Fig. 2a and b, the operating parame-

ters are {0.5, 0.7, 0.8}. For the IGLC, the parameter

set reduces to {l si/ hs, V gi/ V gref }—the values used for

Fig. 2c are {0.1, 0.8}. In the case of the pig-lift, the

motor valve bean is 20/64U, and the only optimization

parameter is {l si/ hs}, set to {0.25} for the elaboration

of Fig. 2d.Except for the pig-lift, the first graph in each set of

plots displays the casing pressure, P cl, at the injection

point (well-head), showing the pressurization up to the

closing of the motor valve (pressure peak) followed by

depressurization while gas flows into the tubing until

the gas-lift valve closes. For the pig-lift, the injection

point is at the top of one of the tubing’s legs and a gas-

lift valve is not employed. The second graph shows the

pressure at the bottom of the well, P wb, which governs

the flow from the reservoir to the well (Eq. (2))—the

lower the average value of P wb, the higher the averagedischarge rate is. The last part of the curve represents

the loading of the slug.

qr ¼ PIð P r À P wbÞ: ð2Þ

The third graph shows the elevation and production

of the slug (and the transfer, in the case of the IGLC

and PL)— hl and hg are the positions of the top and

bottom of the slug, respectively; their difference is the

length of the slug, which diminishes when fallback

occurs.

(d) Description of equation set for each stage of the PL cycle


Injection and transfer Injection valve opens: F gi = 1 [b3,b4.2,b6,b8,b9.1,b11,b15,b16.1,b16.2] F gi = 1, F gp =0, PC = P t3, Hg = H t À hg, Hl = hl, qgt = qgc, v gt = v P.

Elevation Injection leg unloaded: hg = 0 [b3,b4.3,b4.4,b6,b8,b9,b10,b11,b13,b14]

F gi = 1, F gp =0, PA= P gtl, PC = P t3, Hg = H t + hg, Hl = hl, qgt =qcl, qgL1 =qcl, qgL2 =qcl, v gt = v l, F ˜ r = 1, F f = 0.

Production slug reaches the surface: ht = H tl [b3,b4.3,b4.4,b6,b9,b10,b13,b14]

F gi = 1, F gp =0, PA= P gtl, PC = P wh ,, Hg = H t + hg, Hl = ht , qgt =qcl, qgL1=qcl, qgL2= qcl, v gt = v l, F ˜ r = 1, F f = 0.

Decompression gas reaches the surface: hg = H t [b3,b4.3,b4.4,b6,b13,b14]

F gi = 0, F gp =1, PA= P gtl, P t2 = P wh ,, Hg = 2 H t , hg = Ht , qgt = qgL, qgL1 = qLl, qgL2 = qL2, v gt =0.5v gs, F ˜ r = 1, F f = 0.

Loading gas pressure relieved: V gs = 0 [b6,b11,b13,b14]

PA= P t4, PC = P t4, Hl = hlr , F ˜ r = 1, F f = 0.

Cycle restarts Liquid reloaded: hlr = l si/2 * loop back *

(e) General settings

Parameter IGL IGLP IGLC PL

Ach 0 0 Ach 0

Ag At À Af At At À Af At

AG – – Ach At

D – – Dch Dt

Ho H gv H gv H t1 H t PE P t4 PA P t4 PAPF – P c4 P c2Q F gv qgv F gv qgv F gv qgv F gi qgiwP 0 wP 0 0

Table 4 (continued )


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In Table 5, the simulator results for the conventional

IGL are compared to Liao’s calculations, which were

tuned to the data of Brown, Brill and Neely. A satisfac-

tory degree of agreement is observed. Accurate labora-tory data is scarce for the gas-lift with plunger and the

pig-lift. Nevertheless, the order of magnitude of the

elevation time obtained in the simulations is equal to

that observed in the field, though values are only good

as 10%, to the best of the authors’ knowledge. The

velocity of the real plunger and pig would be influ-

enced by the fallback, because the slug weight would

decrease along the way up to the surface. Unfortunate-

ly, laboratory data is even more difficult to find for the

gas-lift with chamber. The order of magnitude of the

production gains, over the CGL and IGL, calculated in

the simulations are in the range that is observed in the

field, to the best of the authors’ knowledge.

3.3. Optimization

The immediate application of the simulator model

is the optimization of the lift operation, i.e., the

maximization of a ‘‘goal’’ by variation of the oper-

ation parameters. This ‘‘objective’’ function might be

the daily production (volume of oil) or daily profit

(difference between economic gain and cost) for

Fig. 2. (a) Sample of the simulator output for the IGL’s example. (b) Sample of the simulator output for the IGLP’s example. (c) Sample of the

simulator output for the IGLC’s example. (d) Sample of the simulator output for the PL’s example.


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instance. Here, the most economical point of oper-

ation—maximum daily profit— is taken up as the

criterion of comparison among the different IGL

methods. For the sake of example, market values in

Brazil, at the time of the development of this work,

were adopted to define the economic conditions,

leading to the constraint expressed by Eq. (3), relating

the increment in gas input, dQgas, to an increase in oil

output, dQoil, for an open system, all factors being

constant but the volume of gas injected.

dQgasdQoil

< 1000: ð3Þ

For the same reservoir case used for Fig. 2, the

most economical point was found to be {0.3, 0.7, 0.5}

for the IGL, {0.3, 0.6, 0.58} for the IGLP, {0.1, 0.49}

for the IGLC and {0.4} for the PL. Further simula-

tions yielded the optimum operating point for the IGL

methods, under various reservoir conditions (PI and

P r ), as shown in Table 6. Each cell of Table 6 gives the

daily liquid production Qlp and daily gas consumption

Fig. 2 (continued ).

Table 5

Comparison of Liao’s model with the present model

Variable Liao’s model Present model

Volume of liquid slug

reaching the surface (m3)

0.309 0.299

Time for elevation (s) 275 289

Average bottom-hole

flowing pressure (MPa)

3.4 3.8


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Qgi, with the parameter settings for the most econom-

ical point.

3.4. Selection

The results presented in the previous section are

useful for the adjustment of the operation, given dif-

ferent work conditions. Another important use of the

simulator model is the selection of the best IGL design.

It must be emphasized that other engineering factors

are also relevant for the final decision, and should be

considered in an integrated analysis.

For the purpose of comparison, once the optimum

operating point was determined for each method,

under different reservoir condition (PI and P r ), then,methods were ranked following the economic crite-

rion given by Eq. (3). The continuous gas-lift (CGL)

was included in the analysis too. The results are

presented in Table 7 with the best option placed at

the top of each cell, where the relative position of the

conventional IGL emphasized in gray, for reference.

In some of the cases shown on Table 7, when the

differences in gain are not too large, other design and

operation issues will determine which method to

choose (Clegg et al., 1993). If the selection criterion

of maximum oil production is adopted, the rankingwill certainly be modified.

For the cases studied, the conventional IGL per-

forms reasonably well, ranging from average to best

economic performance, usually outdoing the CGL.

Although the conventional IGL is not always the best

among the IGL methods, it may still be selected due to

its simplicity and reliability, especially when the gains

offered by the other methods are only marginal. Also,

it was determined that the ideal IGLP performs very

well for low-pressure reservoirs. For high-pressure

Table 6

Economic optimum

hs/ H w, P r (MPa) PI (m3/day MPa)

2.5 5 10 20(d) PL

0.2, 2.78 2.2 4.4 8.4 16.9

965 1750 2500 5312

60, 70, 41 60, 70, 41 60, 70, 50 40, 60, 58

0.4, 5.57 8.6 14.8 27.6 42.5

2518 3032 4886 6328

30, 70, 41 40, 70, 41 40, 70, 50 50, 70, 50

0.6, 8.35 14.0 25.9 48.0 80.0

2521 4136 6773 11,469

30, 70, 41 30, 70, 41 30, 70, 50 30, 60, 58

0.8, 11.13 19.4 36.8 67.9 106.0

2726 4705 8539 9868

30, 70, 41 30, 70, 41 30, 70, 50 40, 70, 41

(b) IGLP

0.2, 2.78 3.1 5.8 10.6 17.9

670 1273 2319 3912

50, 60, 70 50, 60, 70 50, 60, 70 50, 60, 70

0.4, 5.57 9.2 17.2 28.1 47.3

1765 3303 4491 8347

30, 60, 70 30, 60, 70 40, 70, 70 40, 60, 70

0.6, 8.35 14.8 27.7 48.1 70.5

2509 4708 7002 10,052

30, 60, 70 30, 60, 70 30, 60, 58 50, 70, 70

0.8, 11.13 20.2 37.9 67.3 100.0

2736 5122 9026 12,812

30, 60, 58 30, 60, 58 30, 60, 58 40, 65, 60

(c) IGLC

0.2, 2.78 4.4 7.1 13.5 24.6

745 1290 2389 4222

50, 70 30, 53 30, 53 30, 53

0.4, 5.57 9.6 17.8 32.9 54.8

1435 2634 5272 8578

20, 60 20, 60 20, 70 20, 70

0.6, 8.35 16.0 29.5 51.4 83.9

2391 4215 6891 10,252

10, 49 10, 49 10, 49 10, 49

0.8, 11.13 21.9 40.0 68.8 109.0

3107 5447 8804 12,620

10, 60 10, 60 10, 60 10, 60

(d) PL

0.2, 2.78 2.8 5.5 10.1 17.7

1200 2300 4248 7366

70 70 70 70

0.4, 5.57 7.4 13.7 23.9 38.2

1750 3239 5619 8812

60 60 60 60

0.6, 8.35 13.5 24.2 39.2 65.7

3031 5326 8371 16,142

40 40 40 30

Table 6 (continued )

hs/ H w, P r (MPa) PI (m3/day MPa)

2.5 5 10 20

(d) PL0.8, 11.13 21.4 39.5 65.4

5400 9603 14,764

25 25 25

Reading key for the table cells: Qlp (m3/day).

Qgi (m3/day).

l si/ hs, P to/ P co, V gi/ V gref (%, %, %).


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reservoirs with high PI, the ideal IGLP performance

drops below the conventional methods —IGL and

CGL. It must be noted that the conventional IGL

comes a close second to the ideal IGLP. These resultsindicate that reduction of the fallback may not be the

critical factor for the improvement of the performance

of intermittent gas-lift. In general, the IGLC performs

very well for all the reservoirs studied, with the

advantage that its performance does not degrade too

much, comparatively to the other methods, as the

reservoir pressure decreases. Since pressure depletion

will occur along the life of a reservoir, the IGLC

would be a good long-term choice. The IGLC fares

better than its competitors because its liquid produc-

tion compensates its gas consumption, even when thegas input is greater than that of the other methods,

especially the CGL. The installation of the IGLC is

very similar to the IGL, but the IGLC produces more

liquid and is more economical than the IGL. The ideal

pig-lift performs poorly for high-pressure reservoirs,

mainly because of its high gas consumption. For low-

pressure reservoirs, the ideal PL does better than the

CGL, but, again, it falls behind the other methods due

to the gas input, even though its liquid output is as

good as the IGLP and better than the IGL. It must be

noted that the PL installation is somewhat simpler

than the others considered here.

4. Dynamics

The engineer must dominate the subject of his

work, therefore it is not enough to compute the IGL’S

output; the relationships among the variables must be

understood as well. The simulator model is helpful in

assessing the measure of the influence of any operation

control parameter on the dependent variables of the

IGL. The following sections describe some of these

studies providing insight into the IGL’s dynamics.

4.1. Variables

The daily production rate QL is equivalent to the

average discharge per cycle—the ratio of the volume

of liquid produced at the surface per cycle, V Ls, to the

duration of one cycle t 1. These two variables are the

main concern of the present analysis. Some of the

parameters that affect them are: the injected volume of

gas per cycle, V gi; the gas-lift valve’s seat diameter,

d gv; the initial length of the liquid slug (load), l si; the

casing pressure ahead of the gas-lift valve at the time

of its opening, P co.Other important variables condition the behavior of

IGL systems too, but they usually do not offer ample

opportunity for control; as so they are assumed con-

stant in all of the foregoing discussion: (a) geometry

of the well—casing diameter, Dc, tubing diameter, Dt ,

well perforation depth, H w, tubing length, H t ; (b) fluid

properties — specific weights of oil, co, water, cw, and

gas dg, water volumetric fraction, /w, reservoir gas–

liquid ratio, RGL, viscosities of oil, lo, water, lw, and

gas, lg; (c) reservoir state—static pressure, P r , pro-

ductivity index, PI, geothermal gradient, aT; (d) sur-face facilities—well-head pressure, P wh, gas injection

pressure, P gi, surface temperature, T s. For simplicity,

H t was set equal to H w, and the depth of the gas-lift

valve, H gv, was also kept constant when comparing

the IGL methods.

For one type of gas-lift valve, all geometric param-

eters are known once the seat nominal diameter is

given—the diameter of the flow orifice, the open–

close characteristic curve, and the flow rate perform-

ance curve. The initial length of the liquid slug defines

Table 7

Selection rank for IGL methods (economic criteria)


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the load to be elevated, is determinant of the ‘‘counter-

pressure’’ in the tubing, and, specifically, of the tubing

pressure downstream of the gas-lift valve at the mo-

ment of its opening, P to. The load will be represented by the non-dimensional parameter L = l si/ hr , where hr is the height of the column of liquid equivalent to the

reservoir static pressure ( P r / qL g ), or the maximum

load. The opening pressure of the valve on the casing

side will be represented by the non-dimensional para-

meter R = P to/ P co. The closing pressure, P cc, becomes

implicitly defined. The total amount of gas injected

will be presented by non-dimensional parameter

V = V gi/ V gref , where V gref was originally defined in

the work of Liao (1991). V gi also establishes the dura-

tion of injection by the motor valve at the surface. Al-

ternatively, the minimum amount of the gas needed to

trigger the opening of the gas-lift, V g, valve may be

chosen as the reference volume, since its value is un-

ambiguously related to the parameters mentioned

above. The reservoir static head will be presented by

P = P r / P H, where P H = q gH w is the static head of the

well filled up with liquid (note also that P = hr / H w).

Finally, due to a lack of standards, the productivity

index of reference is arbitrarily set to PI* = 1 m3/day

MPa, in the definition of the non-dimensional index

I = PI/PI*. The study was conducted with the case of

Table 2, for 0.2 < P < 0.8 and 2.5 < I < 20, but all thegraphics presented here are for P = 0.6 and I = 10. For

the sake of comparison, the liquid volume fed by the

reservoir to the well, in the first cycle, is used as a

measure of discharge (Santos et al., 1998).

4.2. The conventional IGL

4.2.1. Influence of the injected volume of gas, V gi , on

V Ls , t 1 and Q L

Fig. 3 displays results with d gv = 12.7 mm (1/2 in.),

L = 0.5 and R = 0.7. The simulation shows two inferior limits for V gi: the minimum value to open the valve

(6), and the minimum value for the slug to reach the

surface (5). One of the two becomes the operation

minimum value for each particular installation. Notice

in Fig. 3(1) that V Ls is smaller than V Li, the initial

volume of the liquid slug; that is due to the fallback,

the slug’s loss of liquid during its rising— DV L. Also,

it is observed that part of V Ls is provided by the

produced slug, and the rest comes from a fraction of

the fallback dragged by the gas during the decom-

pression phase. The volume of the produced slug

increases monotonically with V gi, but is non-linear,

and the rate of increase becomes very small for larger

values of V gi. The amount of liquid dragged by the gasis significant only for high values of V gi. These

observations are explained by the relationship between

V gi and the ascent velocity. Fig. 3(2) shows that, for the

same depth, the elevation period diminished as V giincreases; therefore the slug has less time to lose

liquid. Besides, the gas at high speed is more proficient

to drag the liquid film left behind by the slug.

Fig. 3(3) shows the moment that the gas-lift valve

closes (t vc), comparing it to the moments of arrival at

the surface of the top and bottom of the slug (t Le, t Lp),

which, respectively, define the elevation and produc-

tion phases. For low values of V gi, the valve closes

during elevation, thus the lift proceeds sustained by

the gas energy stored in the tubing. For high values of

V gi, the valve remains open during production, con-

tinuously supplying energy for maintenance of the

process, including the dragging of the liquid film,

when the valve stays open after the production phase.

The increase in V Ls, discussed above, does not mean

a gain in production rate, which is also dependent on

the period of the cycle, t 1, which, in turn, is the sum of

the periods of each phase. The period of the elevation

phase diminishes as V gi increases, but it tends to acertain limit (Fig. 3(4)); on the other hand, the decom-

pression per iod t gd À t Lp always increases as V giincreases (Fig. 3(4)), because there is more gas to be

purged from the tubing. This last effect becomes

predominant and t 1 increases. The combination of the

effects of V gi on V Ls and t 1 results in the behavior of the

discharge rate, QL, shown in Fig. 3(5)— for smaller V gi,

the increase in V Ls becomes an increase in QL, while for

larger V gi, the increase in t 1 translates into a decrease of

QL. Therefore, a ‘‘critical’’ value of V gi exists, deter-

mining the maximum daily flow rate (note that it is not at the same point where V Ls is at its maximum).

The maximum discharge is usually associated to

very large gas injection rates, consequently, the costs

may overcome the gains, and the point of maximum

daily profit will be different from the point of max-

imum daily flow rate. For the sake of the practice in the

field, it is important to emphasize the sensitivity of the

discharge rate with regard to V gi — with amplitudes of

the order of 15% to 20% of the median of the range of

QL, when 0.3 < V < 1.5, for the current examples.


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4.2.2. Influence of the initial length of the slug (load),

l si , on V Ls , t 1 and Q L

For a well-reservoir system (given P and I ) with

the R and V parameters set at fixed values, the larger the load L, the larger the volume of the produced

liquid. This happens because the distance to reach the

surface is smaller for the larger l si, while the average

speed does not change as much, if R and V are kept

constant; therefore, the travelling time is also smaller,

and there is less time for the slug to lose liquid during

the elevation phase (see Fig. 3(6), where the fallback is expressed as FB = 1 À V Lf / V Li). Surely, it takes more

time to produce the longer slug, which allows for

more penetration of the gas shot into the body of the

Fig. 3.(1)IGL: L = 0.5, R = 0.7.(2)IGL: L = 0.5, R = 0.7.(3)IGL: L = 0.5, R = 0.7.(4)IGL: L = 0.5, R = 0.7.(5)IGL: L = 0.5, R = 0.7.(6)IGL:V o = minimum

V o = minimum to open the valve. (7) IGL: V o = minimum to open the valve. (8) IGL: L = 0.5, V = V o. (9) IGL: L = 0.5, V = V o. (10) GLI: L = 0.5,

V = V o. (11) IGLC: L = 0.1. (12) IGLC: L = 0.1. (13) IGLC: L = 0.1; IGL: L = 0.5 (same hsi). (14) IGLP: L = 0.5, R = 0.7. (15) IGLP: L = 0.5,

R = 0.7. (16) IGLP: L = 0.5, R = 0.7. (17) Pig-lift.


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liquid, augmenting the fallback during the production

phase. Nevertheless, as the slug is produced, the liquid

mass in the tubing diminishes, causing an acceleration

of the slug that lessens the penetration; effectively, the

augmentation of the fallback during the production

phase is offset by its reduction during the elevation

phase. Apparently, it would be desirable to operate

with larger loads, but, then, the cycle’s period t 1would also increase, mainly due to the longer period

of recharge (Fig. 3(7)). Since the effects of L on V Lsand t 1 tug QL in opposite directions, one must use the

simulator to search for the ‘‘critical’’ value of L, for

which the discharge rate is maximum, or for the

‘‘economic optimum’’ value of L, different from its

‘‘critical’’ value, that maximizes the daily profit. The

sensitivity of QL to L is high—with amplitudes of the

order of 40% of the median of the range of QL, when

0.2 < L < 0.6, for the current examples.



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4.2.3. Influence of the casing pressure at the opening

of the gas-lift valve, P co , on V Ls , t 1 and Q L

In this example, the valve’s seat diameter was set to

d gv = 12.7 mm (1/2 in.), while the load and the injection

were fixed, respectively, at L = 0.5 and V = V o (V o = V gv/

V gref , the minimum value required to open the valve).

Under such conditions, V Ls is greater for the smaller

values of R, i.e., larger P co (Fig. 3(8)). This occurs

because the higher the pressurization of the casing is,

the higher the rate of injection in the tubing will be,

yielding a greater speed of ascent which, in turn, bears a

larger V Ls, as argued back in Section 4.2.1. At the same

time, a larger P co implies more gas to open the gas-lift

valve, V gv, and, as seen before in Section 4.2.1, more

gas means more liquid at the surface. With regard to the

cycle’s period, the higher speed of the slug translates



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into a shorter elevation period, but the larger the

amount of gas requires a longer period for decompres-

sion. The final balance is unfavorable, resulting in an

increase of t 1, as P co increases (decreasing R, Fig. 3(9)).Again, these effects exert contrary pulls on QL, creating

a ‘‘critical’’ point with respect to R, where QL has a

maximum value (Fig. 3(10)). The sensitivity of QL to R

is small, comparatively to the sensitivity associated to V

and L —with amplitudes of the order of 6% of the

median of the range of QL, when 0.5 < R < 0.75, for the

current examples.

4.3. The IGL with chamber

In the case of the IGLC, the operation parameters

studied were V and L. In the following example, a pi-

lot valve was employed with seats of 7.9 mm (5/16 in.)

and 14.3 mm (9/16 in.), for the pilot and main sections,

respectively, and set to open at the pressure available at

the surface compression facility. For the IGLC, l si is

equal to the length of the accumulation chamber, de-

fining the slug load and the period of charge, just as in

the case of the conventional IGL. The difference is in

the height of the slug when lift begins, after the phase

of liquid transfer from the chamber into the tubing: for

the IGL, hsi = l si, while, for the IGLC, hsi = l si.( Dc/ Dt )2.

The casing pressure and the tubing counter pressureare determined by these variables.

The simulations have shown that the behavior of the

IGLC is very similar to that of the IGL, as presented in

the previous sections. The differences lie on the posi-

tions of the ‘‘critical’’ and ‘‘economic optimum’’

points. It is worthy of note that, at these points, the

discharge rate of the IGLC is consistently greater than

that of the IGL, while the values of L are lower, as well

as the fallback. Results are displayed in Fig. 3(11,12

and 13). The advantage obtained by the IGLC comes

from its low average well-bottom pressure.

4.4. The IGL with plunger

The ideal IGLP is a model in which leakage does

not occur around an ideal plunger, hence there is no

fallback — seemingly a desirable feature. Accordingly,

the produced volume of liquid is equal to the load, and

is not affected by V gi, P co or d gv. Still, there is min-

imum required value for V gi, dependent on L and R,

for the slug to reach the surface—for instance, with

d gv = 12.7 mm (1/2 in.), L = 0.5 and R = 0.7, V min= 0.6.

Notice that, under the same conditions, the conven-

tional IGL is able to produce with V = 0.4, consuming

less gas. For V above V min, the ideal IGLP’s V Ls islarger than the IGL’s, due to the absence of the fall-

back, but for even higher values of V , the IGL’s V Lsapproaches the same ideal value of the IGLP (Fig.

3(14)). The larger V Ls of the ideal IGLP is not neces-

sarily an advantage (Santos et al., 1998), because its

cycle’s period is adversely affected. The ideal IGLP

cycle is longer than the IGL cycle (Fig. 3(15)), result-

ing in discharge rates of the same order (Fig. 3(16)).

In some cases, the ‘‘economic optimum’’ belongs to

the IGL because the IGLP consumes more gas. The

IGLP’s cycle is longer because its slug remains larger

(no fallback): first, the slug’s ascension is slower due

to the greater opposition of friction (more length) and

gravity (more weight), resulting in a longer elevation

period; second, the production period is longer be-

cause the slug’s speed is lower and the slug’s length is

longer; third, the loading (recharging) phase is also

longer because only the reservoir contributes to it,

while, in the IGL case, part of the fallback returns to

the bottom of the well, completing the load sooner.

4.5. The IGL with pig

The following observations regard the ideal pig-

lift, where the ideal foam-pig does not allow any

leakage. Consequently, for the same load L, V Ls is

invariable as in the IGLP case. Besides, there is not a

gas-lift valve, so the only parameter left is the length

of the slug. Evidently, V Ls = V li, which is proportional

to L — there is no need for simulation here. But, as L

increases, so does the cycle’s period t 1 (Fig. 3(17)—

left), leading to lower discharge rates (Fig. 3(17)—

right). To shorten the cycle’s period, both legs of the

tubing are connected to the discharge port for simul-taneous exhaust of the gas thus reducing the decom-

pression period (Fig. 3(17), dotted lines).

5. Conclusion

This paper presents a study of the influence of the

operation parameters of intermittent gas-lift systems

on the performance of four alternative methods, car-

ried out through a new consistent full-transient simu-


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lation model, developed for this task and disclosed for

the public domain. Due to space limitations, only the

highlights of the study are discussed, and a few

examples are displayed. Nevertheless, important rela-

tionships between the variables of the process are

clarified, providing insight into the inner workings of

the IGL system. Running more cases with the simu-

lator, the engineer may further his understanding of

the system. The results show the sensitivity of the IGL

to the parameters of its design. Therefore, the selec-

tion of the best method, whatever criteria are consid-

ered, requires a careful analysis, as well as the

determination of each method’s optimum point of operation. The simulator is necessary to resolve these

complex tasks. A comparative study of the methods is

also presented. Again, the simulator proved to be an

invaluable tool. Although limited in scope, the study

reveals that ranking the IGL methods is not a trivial

job, and also challenges some unfounded myths. The

examples show the potential of the simulator to build

comparison tables that may be an aid for ballpark

assessment of the methods suitability and optimum

settings of the operating parameters; then, the simu-

lator can be used again to refine the project. Thesimulator may also be coupled to more sophisticated

computer-aided design and selection programs, as one

of the CAD&S’s processing modules.

Work is under way to improve details of the si-

mulator—either its physical or numerical models, but

what are the most needed, today, are carefully con-

trolled experimental works, with large-scale fully ins-

trumental apparatuses, and more accurate and reliable

field data. A laboratory rig will be designed to provide

a reliable database for a definitive validation of the si-

mulator. As it stands, numerical values appearing in

this article must be viewed with prudence; neverthe-

less, the qualitative observations on the variables’

relationships and the comparisons among the IGL

methods may still hold, since the same model assump-

tions and equations were applied to all methods. As the

work progresses, the authors pledge to keep the pet-

roleum engineering community up to date with further

developments.

In 1993, gas-lift installations reached over half of

about 100,000 wells in the USA, each delivering at

least 540 m3/year (3400 bbl/year), and accounted for

the majority of offshore wells (Clegg et al., 1993)Most of these gas-lift devices employed the CGL

method. Although few official data is available, it is

quite certain that gas-lift is largely used in Brazil also

(Fig. 4 illustrates two examples), and will continue to

be, as a good alternative for reservoirs of low pressure

or low productivity. Hence, proper selection and

optimization of IGLs are of paramount importance.

Acknowledgements

The authors wish to express their thanks to Mr.Attilio Triggia, MSc for valuable suggestions. Thanks

are also due to PETROBRAS, CEPETRO/UNICAMP

and the Dept. of Petroleum Engineering of FEM/UNI-

CAMP.

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Fig. 4. Examples of lift energy and method share by number of wells, year 2000 (private compilation). RESERV = reservoir own energy;

MECH= external mechanical energy; GAS = external gas energy. Relative sizes of parts represent percentages of the whole.


7/16/2019 Gas Lift


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