REVIEW AND APPLICATION OF THE TULSA LIQUID JET PUMP MODEL

pump

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REVIEW AND APPLICATION OF THE TULSA LIQUID JET PUMP MODEL

Pål Jåtun Pedersen

Trondheim December 2006

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Review and Application of the Tulsa Liquid Jet Pump Model December 2006

Preface

This report is a mandatory project assignment in the 9th semester of the petroleum production engineering studies at NTNU It was written at the institute for petroleum technology and applied geophysics, fall 2006 The assignment consists of 71 pages, and was delivered the 19th

of December 2006

I would like to thank Professor Jón Steinar Guðmundsson for good help and advice

throughout the project Also, I am very grateful for all the help I have got from the people at Petroleum Experts Ltd., regarding the version update of PROSPER

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Summary

As a water drive reservoir is depleted, production will fall and inflow decrease As a

consequence of this, the well will either stop flowing or produce only a limited amount of oil and gas In these cases, an artificial lift system can be installed to increase the production and save the well One of these lift systems is the Jet Pump

The Jet Pump operates on the principle of the venturi tube, converting pressure into velocity head by injecting power fluid through a nozzle This creates a suction effect which drives the production fluid through the pump At the diffuser the velocity head is converted into

pressure, allowing the mix of power and production fluid to flow to the surface through the return conduit

dependent on these values Especially is it dependent on the throat-diffuser friction factor, which again depends on the gas-oil ratio

Calculations were performed on a North Sea well with a gas-oil ratio on 95 Sm3 Sm3, using both the Tulsa model and models based on incompressible flow The calculated efficiency was, as expected, higher for the models based on incompressible flow

The well performance program PROSPER was used for pressure drop calculations Also, the Jet Pump function in the program gave about similar results as the Tulsa model Perhaps is the Tulsa model used as the Jet Pump function in this program Anyhow, the similarity in results between the Tulsa model and PROSPER indicates that the calculations performed in this project is reasonable and that the model is applicable to a field situation as presented here

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Table of contents

1 Introduction 1

2 Jet Pump Literature Survey 2

2.1 When is artificial lift required 2

2.2 Jet Pump compared to other artificial lift methods 3

2.4 Jet Pump principles 4

3 Review of the Tulsa Jet Pump Model 5

3.1 Development of the model 5

3.2 Presentation of the model and its main principles 6

4 Tulsa Jet Pump performance 9

4.1 Main factors to control pump performance 9

4.2 Sizing of the pump 10

5 Application of the Tulsa Jet Pump Model 13

5.1 Sizing and performance calculations 13

5.2 Evaluation of results 17

6 Application of the Tulsa model on a North Sea well 18

6.1 Case description 18

6.2 Model calculations 19

6.3 Evaluation of results 22

7 Application of Other Models on a North Sea Well 23

7.1 JSG model calculations 23

7.2 NTNU project calculations 24

8 Discussion 28

9 Conclusion 30

10 References 31

11 Tables 32

12 Figures 34

13 Appendixes 46

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Several different Jet Pump models have been developed, varying in accuracy and complexity However, few models for predicting the behaviour of compressible flow are developed Among these few models is the “Performance model for Hydraulic Jet Pumping of two-phase fluids” by Baohua Jiao, published in a thesis at Tulsa University in 1988 (in this project referred to as the Tulsa model)

The project assignment is to review the Tulsa model, convert the basic equations to SI and perform calculations for a production well in the North Sea, using PROSPER for pressure drop calculations Then, look at previous NTNU projects/thesis and perform calculations on the North Sea well with a few other models Finally, compare the results with the Tulsa

model

The project starts with a literature survey of the Jet Pump, giving a brief introduction to the Jet Pump principles, different Jet Pump models and comparison between Jet Pump and other artificial Jet Pump models In chapter 3, 4 and 5 the Tulsa Jet Pump model is introduced and documented, and in chapter 6 the model is used for calculations on a North Sea well Chapter

7 contains calculations using other models, for comparison

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2 Jet Pump Literature Survey

2.1 When is artificial lift required

The objective of any artificial lift system is to add energy to the produced fluids, either to accelerate or to enable production

Some wells may simply flow more efficiently on artificial lift, others require artificial lift to get started and will then proceed to flow on natural lift, others yet may not flow at all on natural flow In any of these cases, the cost of the artificial lift system must be compared to the gained production and increased income In clear cut cases, such as on-shore stripper wells where the bulk of the operating costs are the lifting costs, the problem is usually not present In more complex situations, which are common in the North Sea, designing and optimising an artificial lift system can be a comprehensive and difficult exercise This

requires the involvement of a number of parties, from sub-surface engineering to production operations

The requirement for artificial lift systems are usually presented later in a field’s life, when reservoir pressure decline and well productivity drop If a situation is anticipated where

artificial lift will be required or will be cost effective later in a field’s life, it may be

advantageous to install the artificial lift equipment up front and use it to accelerate production throughout the field’s life

All reservoirs contain energy in the form of pressure, in the compressed fluid itself and in the rock, due to the overburden Pressure can be artificially maintained or enhanced by injecting gas or water into the reservoir This is commonly known as pressure maintenance Artificial lift systems distinguish themselves from pressure maintenance by adding energy to the

produced fluids in the well; the energy is not transferred to the reservoir (Jahn, Cook & Graham, 1998)2

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2.2 Jet Pump compared to other artificial lift methods

The Jet Pump has many advantages towards other artificial lift systems There are no moving parts, the pump is tolerant not only of corrosive and abrasive well fluids, but also of various power fluids Maintenance and repair are infrequent and inexpensive, the pump can be

replaced without pulling the tubing (casing type installation) and it consists of few parts The pumps are suitable for deep wells, directional wells, crooked wells, subsea production wells, wells with high viscosity, high paraffin, high sand content, and particularly for wells with GOR up to 180 3

The casing type installation is the most common solution, using the casing-tubing annulus as the return conduit and the tubing as the power fluid string For this type of installation, the production of free gas through the pump causes reduction in the ability to handle liquids The advantage is, as mentioned above, that the Jet Pump can be replaced without pulling the tubing (Brown, 1982, Jiao, 1988)1,3

Figure 4 shows a comparison for the different artificial lift methods

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2.4 Jet Pump principles

Jet Pumps operate on the principle of the venturi tube A high-pressure driving fluid (“power fluid”) is ejected through a nozzle, where pressure is converted to velocity head The high velocity – low pressure jet flow draws the production fluid into the pump throat where both fluid mix A diffuser then converts the kinetic energy of the mixture into pressure, allowing the mixed fluids to flow to the surface through the return conduit (Jiao, 1988)1

Figure 2 illustrates the principle

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3 Review of the Tulsa Jet Pump Model

3.1 Development of the model

Experimental studies were performed using a mixture of water and air as the production fluid and water as the power fluid The operating pressures were set to typical values found in the field, with power fluid, for example, reaching 3000 psig (20 MPa) and production intake fluid exceeding 1200 psig (8.3 MPa) The performance data acquired were the power fluid

pressure, the pressures at the intake and discharge, the flow rates of the power fluid, the two phases of the production fluid, and the appropriate temperature so that the air-liquid ratio could be computed For further description of the experimental facility and test data it is referred to the thesis

The analysis of the data followed the model of Petrie, Wilson and Smart (PWS) This model

is based on conservation of mass and energy, and is widely familiar to production engineers The PWS model and the Tulsa model differ only in the treatment of the two empirical,

dimensionless parameters, and , which are the loss parameters for the nozzle and the throat-diffuser, respectively The objective of both models is to predict a dimensionless

pressure recovery ratio, N, as a function of a dimensionless mass flow ratio, M (Jiao, 1988)

n

K K td

1

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3.2 Presentation of the model and its main principles

The model is originally derived in oilfield-units, but is presented here in SI-units Conversion from field to SI-units is a task specifically mentioned in the project-description, and is

conducted on both the “Derivation of the Jet Pump Model” (Appendix A) and the “Pump Sizing Procedure” (Appendix B) Following is a presentation of the main principles of the Tulsa model For the model derivation in its entirety it is referred to Appendix A The

terminology used in the model is detailed in the Nomenclature (Appendix A, page 61-62) and shown in Figure 1 The brackets on the right side of the mathematical expressions contain the equation number in the derivation

As mentioned earlier, the purpose of the model is to predict pressure recovery, N, as a

function of dimensionless mass flow ratio, M

The dimensionless pressure recovery is the pressure increase over the pump divided by the pressure difference between the drive fluid and the pump discharge Mathematically it’s defined as follows:

ake

Q

Q Q

Q m

for one phase flow, assuming equal density for the two fluids

Extended to include gas, the mass flow ratio can be expressed as:

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As shown in the derivation of the pump model (Appendix A).The numerator in equation (37) describes the total producing fluid mass flow This includes both liquid and gas, where the term 1.227×Q iarepresents the gas mass flow (derivation page 56-58)

In the Tulsa thesis, it is assumed equal density for the power fluid and the produced liquid phase For an oil production case with high water cut, it could be argued that in equation (37) should be adjusted for difference in oil and water density

=

The model use a functional form of N = f (M)that is based on work by Cunningham4, who developed this function on mass energy conservation principles Simplifying the typing of this function, two component elements are defined:

)1/(

))(

21

(

2 2

)1

K

C K B

N

td n

td

)1()

1

(

)1(

++

−+

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In the above expression of N, the importance of the loss parameters is obvious The nozzle loss parameter, , is in this model set to 0.04 This value was estimated in the Tulsa thesis from optimization based on high pressure data

33 0 63 0 33

2 3

)())(

10

*88.10

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4 Tulsa Jet Pump Performance

4.1 Main factors to control pump performance

The performance of the jet pump can be expressed by comparing different important elements

of the Jet Pump model Figures 7-16 are based on data from the thesis-experiment described

in Chapter 3.1 “Development of the Model” They illustrate jet pump performance under varying conditions

Figure 7 is a plot of the throat-nozzle loss parameter versus air-water-ratio, for three fixed values of (ratio of discharge pressure to power fluid pressure) The trend shows that an increasing air-water ratio results in an increasing throat-nozzle friction factor Figure 8 are the same plot as Figure 7, but with AWR in field units is also expressed in figure 9 Here the throat-nozzle friction factor is plotted against for five different values of AWR Clearly it shows that decrease as increase Hence, referring to equation (19) in chapter 3.2, the higher pressure recovery ratio the lower the friction loss in the throat and diffuser Following,

as N increase and M remains the same, the efficiency will increase Also, the horsepower needed to drive the power fluid will decrease as the horsepower requirement varies with N (Appendix B, step 27, 24, 22)

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results in decreasing flow of the production fluid to the power fluid and overall lower pump efficiency The nozzle/throat relation is described in more detail in chapter 4.2

Figure 13 restates the influence of air (gas) in the system The higher AWR, the lower the total efficiency(N*M) and production flow rate to fluid flow rate This can also be seen on figure 14 The last two figures, 15 and 16 describes N vs M and efficiency vs M for different values of influence which again influence N Decreasing leads to decreasing efficiency and decreasing production fluid flow to power fluid flow

p

4.2 Sizing of the pump

Dimensioning a jet pump is an important part of a jet pump installation process The

nozzle/throat combination defines the degree of pump optimization and performance, another consideration is that a minimum area of throat annulus is required to avoid cavitation

Following is a description of these two important elements of Jet Pump sizing:

The nozzle/throat relation

Jet Pump performance is well specific and careful selection of the nozzle/throat combination

is therefore necessary to ensure optimum well performance Due to this fact, manufacturers of Jet Pumps have made a wide range of nozzles and throats available (Figure 5), where the optimum combination represents a compromise between maximum oil production and

minimum power fluid rates

In general, the areas of nozzles and throats increase in geometric progression Because of this, fixed area ratios between nozzles and throats, R, can be established The different

configurations of the nozzle/throat relation are given in Figure 5 A given nozzle (N) matched

to the same number throat (N) will always give the same area ratio, R This is referred to as an

A ratio For a given nozzle(N): B, C, D….ratios represent throats with number N+1, N+2 and N+3 respectively It is possible to match a given nozzle with a throat which is one size

smaller; this is a Acombination (by some manufacturers also referred to as an X

combination) Because of geometric considerations, application of successively smaller

throats is not suitable

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A specific nozzle/throat combination is defined by a number, which refers to the nozzle size, followed by a character which defines the throat size For example a 10A combination refers

to a 10/10 nozzle/throat combination, a 12B a 12/13 combination and so on (Figure 3)

The A(X)-ratio is for high lift and low production rates compared with the power fluid rate, while for instance the C ratio is for low lift and high relative production rates This is

explained in the paper “Jet Pumping Oil Wells” by Petrie, Wilson and Smart:

“Physical nozzle and throat sizes determine flow rates while the ratio of their flow areas determines the trade off between produced head and flow rate For example, if a throat is selected such that the area of the nozzle is 60% of the throat area, a relatively high head, low flow pump will result There is a comparatively small area around the jet for well fluids to enter, leading to low production rates compared to the power fluid rate, and with the energy of the nozzle being transferred to a small amount of production, high heads will be developed Such a pump is suited to deep wells with high lifts

Conversely, if a throat is selected such that the area of the nozzle is only 20% of the throat area, more production flow is possible But since the nozzle energy is being transferred to a large amount of production compared to the power fluid rate, lower heads will be developed Shallow wells with low lifts are candidates for such a pump“ (Petrie, Wilson Smart, 1983, Allan, Moore, Adair, 1989)5,6

Cavitation and sizing of throat entrance area

When sizing a hydraulic Jet Pump for multiphase flow, one of the most important factors is to avoid cavitation

Cavitation can damage the Jet Pump, and the throat in particular When oil reaches the bubble point, it is saturated with gas, so any lowering of pressure means that more gas will come out

of the solution The cavitation phenomenon is caused by the collapse of these gas bubbles on the throat surface as the pressure increases along the jet pump axis (Figure 6) This collapse of vapour bubbles may cause erosion known as cavitation damage and will decrease the jet pump performance

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Within the throat, pressure must remain above liquid-vapour pressure to prevent throat

cavitation damage Note that pressure drops below pump-intake pressure as produced fluids accelerate into the throat mixing zone If pressure drops below the liquid-vapour pressure, vapour bubbles will form The throat entrance pressure is controlled by the velocity of the produced fluid passing through it From fluid mechanics we have the Bernoulli equation that states that as the fluid velocity increase, the fluid pressure will decrease and vica verca

In order to maintain the throat entrance pressure above the liquid-vapour pressure, the nozzle and throat combination must be carefully selected The nozzle and throat flow areas define an annular flow passage at the throat entrance This area decides the velocity of the fluid, and therefore the fluid pressure The smaller flow area, the higher velocity of the fluid The static pressure of the fluid drops as the square of the velocity increase and will reach the vapour pressure of the fluid at high velocities This low pressure can cause cavitation Thus, for a given production flow rate and a given pump intake pressure, there will be a minimum

annular flow area required to avoid cavitation (Grupping, Coppes, 1988, Christ, Petrie,

1989,Petrie, Wilson, Smart, 1983)8,7,5

A step-by-step guide for sizing hydraulic Jet Pumps is enclosed in Appendix B The

procedure was first presented in the Tulsa thesis, and is in this project converted to SI-units

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5 Application of the Tulsa Jet Pump Model

5.1 Sizing and performance calculations

Following is a calculation example for the Jet Pump Model This example is originally

presented in the Tulsa thesis, and is included in this project to illustrate the application of the model Input data has been converted to SI-units, and calculations have been carried out following the “step by step”-procedure enclosed in Appendix B

The two set of calculations (in this project, 5.1 and in the example in the Tulsa thesis) only differs in the computing of the Reynolds number In this project, a considerable higher

Reynolds-number was computed, which results in turbulent flow in the power fluid tubing In the Tulsa thesis, laminar flow was calculated in the power tubing In this project the relation

In both cases, the following criteria for flow-regime determination are used: Reynolds

numbers above 2100 implies turbulent flow (transient between 2100 and about 4000) and below 2100 implies laminar flow This results in different flow regimes for the two cases, as the Reynolds number is different

In the Tulsa thesis, the following calculations are made:

Velocity in power fluid tubing = 4.152 ft/sec

Density of the power fluid = 52.93 lbm / ft3

Diameter of the tubing = 1.995 inches = 0.16625 ft

Viscosity of the power fluid = 5 cp

1.9065

16625.0152.493.52124124

906.1 is the value used in the Tulsa thesis

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The above equation assumes that the diameter is in inches, not in feet This is stated on page

61 of the Tulsa thesis As seen in the above calculations, feet are used as the diameter unit This results in an incorrect Reynolds number

Using inches as the tubing inside diameter unit, we get:

108735

995.1152.493.52124124

μ

ρud

N

This is about the same number calculated in this project, therefore it seems like the

calculations using the standard relation

μ

ρud

=

Re are correct

A conversion of the input data, from field to SI units, are given in Table 1

The example-well from the Tulsa thesis is hereafter called Well A

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Review and Application of the Tulsa Liquid Jet Pump Model December 2006 Review and Application of the Tulsa Liquid Jet Pump Model December 2006

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1616

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Reynolds-number mentioned earlier has very little influence on the pump efficiency and power demand

≅

However, the above calculations are not performed with the optimal nozzle/throat

configuration This was found to be 6D (National, Figure 5) which gives efficiency of 23.7% and power requirement on 82.4 HP

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6 Application of the Tulsa Model on a North Sea Well

As presented in the introduction, one of the tasks in this project was to apply the Tulsa model

to a North Sea oil well and suggest if it is suitable for use in the given well Well 34/10-C-16, located in the “Gullfaks”-field, was chosen as the basis for the North Sea Well Well data were found in the “Gullfaks database” on the IPT computer network For the following case, the reservoir pressure was adjusted such that the well became a candidate for artificial lift The depth of the well was slightly increased and the well was made vertical to simplify the case PROSPER, a well performance, design and optimisation program, was used to make the inflow and outflow curves The Characteristics of the well are found in Table 2 The well is hereafter called Well B

For power fluid, processed oil with the same characteristics as the producing oil is selected (for power fluid oil, GOR=0) The pump supplier is “National” (Figure 5) Surface pump pressure together with pump efficiency for the given well are to be computed In Chapter 6.2 the Tulsa Jet Pump Model is used for these calculations

Inflow and tubing performance curves for the well are found in Figure 18 Data for Well B are found in Table 2 and data for the performance curves are found in Table 3

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6.2 Model calculations

The following calculations were performed using different nozzle/throat combinations The efficiency/power relations for the different combinations are found in Figure 19 Following is the calculations for the optimal combination found, 14D:

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2020

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The annular friction loss for the returning fluid was calculated using PROSPER (table 4), by setting the return conduit to “casing-tubing annulus” and the return water cut to 20% (step 13), PROSPER calculates an annular friction pressure loss of 38 bar for the return fluid flow equal to 1769 sm3/sm3 (step 11)

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The values for power fluid rate, surface pump operating pressure, nozzle friction factor and throat-diffuser friction factor from the calculations in 6.2 were inserted in the Jet Pump

function in PROSPER Based on this input data, the program computed the plot seen in figure

20 PROSPER illustrates the lift by plotting the pump discharge pressure vs liquid flow rate The Jet Pump Model in PROSPER gave a liquid flow rate of 957 Sm3 d Hence, the

difference between the Tulsa Jet Pump Model and the model used by PROSPER is about 4.3% Probably a large part of the difference can be explained with the fact that the

bottomhole pressure is adjusted from 271 to 261 bar when using the jet-pump feature in PROSPER The above calculations uses 271 bar as bottomhole pressure Setting the

bottomhole pressure manually to 271 bar after applying the Jet Pump in PROSPER, the

program gives production flow rate of about 1000 Sm3 dwith the given input data

Probably is the Jet Pump function in PROSPER based on the Tulsa model or a very similar model Anyhow, the program indicates that the Tulsa Jet Pump model gives, in this case, reasonable results

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7 Application of Other Models on a North Sea Well

For a basis of comparison, the following chapters include some other models applied to the case presented in Chapter 6.1

7.1 JSG model calculations

This model is presented in the paper “Modell for strålepumpe” by Jón Steinar Guðmundsson (01.06.06) The JSG model is a simple jet pump model, derived from the basic Bernoulli equation for steady state incompressible flow:

.2

1 2

Const

u

P + =

ρ , and the mass flow relation m G +m L =m M

Neglecting the kinetic energy (very small), we have

L L G

ρ is the density of the power fluid, in this case oil with density 843 kg/Sm3

The Pump discharge pressure ( ) we have from point 16 of the calculations in Chapter 6.2 This can also be used here, independent of the models Pump discharge pressure is 318 bar

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(

56.0843

908)58.01(58.0

58.0)

1

(

58,05.75.10

5.1024

*60

1

*769

*843

24

*60

1

*1000

*908

Sm kg

kg m

m

x

L G

M

L G

=

−+

=

−+

=+

=

ρααρ

ρ

ρρα

efficiency becomes:

%29289.0843

)318503(

908

271880

in Chapter 8

7.2 NTNU project calculations

In the NTNU project “Ejektorpumpe” (Moxnes, 2005)9, a Jet Pump model is presented The model is based on a note by Harald Asheim from 200410, containing some small

modifications In this model, calculations are performed for each part of the pump, divided into inflow, ejection, throat (mixing chamber) and outflow

The Bernoulli based model assumes incompressible flow Following is a presentation of how

to apply the model:

1) Fluid velocity in each part of the pump is calculated: tubing power fluid velocity, ejection power fluid velocity, producing fluid velocity, production fluid suction

velocity, mixed fluid velocity in throat and discharge (outflow) velocity of mixed fluid Using V =Q A

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2) Calculating power fluid ejection pressure:

)(

2

1 2 2

prod suction

prod prod

([

1

suction throat

prod prod throat

e power power throat

e

A P

4) Discharge (outflow) pressure is calculated:

)(

2

1

arg 2 2

e disch throat

throat throat

Where ρthroatis assumed to be homogenous

5) Pressure at nozzle inflow:

)(

e disch prod d

2

12

1( + ρ 2 arg − − ρ 2

=

and

power e disch power e

disch power

power power

2

12

1( + ρ 2 − arg − ρ 2 arg

=

7) Power requirement of surface pump:

p surfacepum

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ompared to the efficiency value calculated in Chapter 6.2, the efficiency here is remarkably

he same calculations were performed on well A from Chapter 5.1, using input data from

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The efficiency for well A is also high compared to the values calculated with the Tulsa model This is further discussed in the next Chapter

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8 Discussion

ter 6 and 7, calculations using the different models gave deviant results

he input data were similar for all three models: well data from table 2 and a preset

Tulsa thesis uses the nozzle and throat-diffuser friction factors to calculate pressure loss

For well B, the NTNU project model calculated a discharge pressure on 332 bar and a

re

required surface pump pressure on 335 bar for a preset discharge pressure on 318 bar I

words, the pressure loss over the Tulsa Jet Pump is higher than over the pump in the other model Calculations performed on well A with both models show the same trend as mentioned above, the Tulsa model computes lower efficiency and higher power requirement

In the NTNU project calculations, the velocity of the mixed fluids at the throat exit

1

mach 0.3 (≈100 m/s) results in compressible flow, which violates the assumption for the Bernoulli equation The velocity decreases as the pressure increases throughout the throat (Figure 3.1 This means that if the velocity is about 100 m/s at the throat-exit, the flow wi

be compressible in the entire throat-section

The NTNU project model is not adjusted to in

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The friction factors in the Tulsa model are based on data from experiments using water and air as the two fluid phases The experiments were conducted under typical field conditions,

luid conditions are calculated The model computes efficiency

r a pump, given the production fluid, power fluid and discharge fluid pressure and density

Tulsa model seems to give very good results It possible that the PROSPER Jet Pump function is based on the Tulsa model presented here,

however it is not certain that these friction factor values are accurate when handling other fluids Further development of the model should probably include experiments with other fluids and fluid properties

In the JSG model, none of the f

fo

The model does not take the difference in power and production volume fluid flow into

consideration It should, however, give quite good results for cases with incompressible fluidsand low power/production fluid flow ratio

Comparing to the results from PROSPER, the

is

or the improved Tulsa model from 199811, anyhow, the similarity in results shows that the model is applicable in the field as it is presented in this project

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9 Conclusion

From the calculations conducted in this project, it is clear that there is a difference between

p models when pumping a two-phase fluid The Tulsa model is

presented models are assuming incompressible

from PROSPER shows that the calculations done here are reasonable Based on this, it

s like the model is applicable to a field situation as it is presented in this project For the

e two phases Experimenting with other fluids and using unequal power and production

er

the presented Jet Pum

extended to include gas, while the two other

fluid could contribute to further development of the model An improved version of the Tulsa model has already been made, and is presented in the SPE Journal in September 199811 Among the improvements are more precise values for both the nozzle and the throat-diffusloss factors

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10 References

rmance model for hydraulic jet pumping of two phase fluids”, (Tulsa,

2 Jahn, F., Cook, M., Graham, M.: “Hydrocarbon exploration and production”,

3 Overview of Artificial Lift Systems”, (SPE Article, 1982)

Fluid

rld Oil,

6 : “Design and Application of an Integral Jet Pump/Safety valve in

7 , H.L.:”Obtaining Low Bottomhole Pressure in Deep Wells With

Allan, Moore, Adair

a North Sea oilfield”,( SPE article, 1989)

Christ, F.C, Petrie

Hydraulic Jet Pumps”, (SPE Production Engineering, August 1989)

Grupping, A.W Coppes, J.L.R., Groot, J.G

(SPE Production Engineering, February 1988)

Moxness, V.W: “Ejektorpumpe”, (NTNU Project, 2005)

10 Asheim, Harald: “Ejektorpumpe”(Note),( NTNU, 2004)

Noronha, F.A.F., Franca, F.A., Alhanati, F.J.S.:

Hydraulic Jet Pumps”, (SPE article, September 1998)

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11 Tables

able 1: Data for Well A, Tulsa thesis example (Jiao, 1988)1

T

Table 2: Well B data for the case presented in Chapter 6 (http://gullfaks.ipt.ntnu.no)

Flowing bottomhole pressure, Pi Pa 27100000

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Table 3: Data table for the performance curves in Figure 18 (PROSPER generated table,

using data from the gullfaks database: http://gullfaks.ipt.ntnu.no)

Table 4: Data table for the return annular liquid flow This is calculated by PROSPER, setting

the return conduit to “casing-tubing annulus” and the return water cut to 20 (step 13, Chapter 6.2) Highlighted are the return liquid flow rate (step 11, Chapter 6.2) and the corresponding friction pressure loss

33

Tiêu đề	Review and Application of the Tulsa Liquid Jet Pump Model
Tác giả	Pồl Jồtun Pedersen
Người hướng dẫn	Professor Jún Steinar Guðmundsson
Trường học	Norwegian University of Science and Technology
Chuyên ngành	Petroleum Production Engineering
Thể loại	báo cáo
Năm xuất bản	2006
Thành phố	Trondheim

Định dạng
Số trang	75
Dung lượng	1,57 MB