The criteria on which a gas-condensate system is based are usually a minimum receipt pres-sure at the downstream terminal, the available (or platform) pressure and the produc-tion profile.
A detailed fluid composition and topography is also required to ensure that reliable results can be obtained using the design methods detailed in this section.
The following is a step by step approach to the typical calculations and considerations required in the design of a gas-condensate system.
4.2.1 Material Selection
The maximum for any line size in a gas-condensate system is often dictated by erosion rather than pressure constraints. As the erosional velocity limit is dependent on the pipe mate-rial, material selection should be conducted at an early stage. It is based on the fraction of acid gases present as well as the operating pressures and temperatures of the flowline. Typically gas with 0.2-2.0 per cent CO, at moderate pressures (300-l 500 psia) would be transported via carbon steel flowlines with inhibition and a corrosion allowance. Significantly higher CO, content or pressures may require the use of corrosion resistant alloys.
The Materials Engineering Group in the company’s Shared Resource at is responsible for material selection.
4.2.2 Pipeline Sizing
Once the material has been selected the minimum pipeline diameter can be calculated for the given production rates and pressure constraints. This work should be conducted using the Mechanistic Model within the BPX MULTIFLO program. This model is available from Version 11 of MULTIFLO and is described in detail in Section 4.3.
If the turndown in flowrates over the life of the line is large it may be necessary to constrain the production profile in order to reduce the diameter of the flowline. This is because the liquid volume that can accumulate in the system increases rapidly at low throughputs, leading to the necessity for frequent sphering or the requirement for large slug catchers.
4.2.3 Erosion Limits
Once a diameter and material have been selected the maximum mixture velocity in the must be calculated in order to check that its value does not exceed the erosional velocity limit. The mixture velocity calculation is conducted by the BPX MULTIFLO program.
The method used to determine the erosional velocity limit is detailed in Section It based on a relationship given in for droplet erosion. If sand is present, particulate erosion may be of greater significance than droplet erosion. Work, sponsored by BP, has been undertaken at Harwell Laboratories to determine the velocity limits for particulate erosion.
The conclusions of this work, and a general discussion on erosion issues, is presented in the BP Erosion Guidlines documents, see Section 3, ref (10)
If the erosional velocity limit is exceeded either the line diameter must be increased or the
Section 4. Gas/Condensate BP Multiphase Design Manual
production profile constrained to reduce the maximum mixture velocity. Alternatively a higher grade of pipeline material could be used duplex stainless steel has a higher resistance to erosion than mild steel (Section 3.1.4).
As well as the erosional velocity limit, also recommends that a line velocity of 60 should not be exceeded to ensure that the level of noise emitted by the flow is not excessive.
4.2.4 Slug Catcher Sizing
Once the material and diameter have been chosen, the required capacity of the catcher at the reception facility must be determined. This is based primarily on the volume of liquid that is produced during sphering of the flowline. This sphered liquid volume is dependent on whether or not the line has reached equilibrium ie. whether, for the given set of operating conditions, the total liquid in the line has reached its maximum value.
For non-equilibrium conditions the total liquid volume produced during sphering is assumed to be equal to the liquid that has entered the line since the launch of the previous sphere. This is dependent on the time interval since the last sphere was
launched, the gas and the liquid loading:
prod (non-equilibrium conditions)
t time interval since previous pig launch (hrs)
The liquid loading is the ratio of liquid to gas and varies with pressure, temperature and compo-sition. For this calculation it is normally calculated for the outlet conditions.
For equilibrium conditions the total liquid hold-up in the is calculated using the Mechanistic Model in The volume of the liquid slug produced during sphering is less than the total equilibrium value as liquid flows out of the line during the passage of the sphere. The calculation procedure for determining the sphered slug volume is given in Section 4.4.
The approximate time period required to reach equilibrium conditions is given by:
t (non-equilibrium conditions)
where:
time interval to reach equilibrium (hrs)
BP Multiphase Design Manual Section 4. Gas/Condensates
equilibrium liquid volume (bbls) gas (mmscfd)
4 liquid loading
The sphered liquid volumes for a given system under both the equilibrium and equilibrium conditions can be plotted against gas on a design diagram:
600
This figure shows that for equilibrium conditions the sphered liquid volume decreases with gas flowrate, whereas for non-equilibrium conditions, and a set sphering interval, the liquid volume increases with flowrate. The maximum sphered volume can therefore be determined at the intersection of the sphered liquid volume lines for the equilibrium and non-equilibrium condi-tions. This is often used as a design basis for determining the required working volume of the slug catcher. The location of this intersection moves to a higher flowrate, and therefore a smaller value of the maximum sphered volume, as the sphering frequency is increased.
Typically the intersection between the equilibrium and non-equilibrium lines for sphering interval of 24 hours will occur at approximately 50 per cent of the design throughput. If the intersection occurs at a higher the line is probably oversized. However, this may be unavoidable given the imposed constraints of production profile and allowable pressure drop. At a
greater the value at the intersection, sphering is not necessary to limit the line’s liquid inventory.
At flowrates less than that at intersection the maximum time interval between spheres can be calculated from:
t (non-equilibrium conditions) Q 4
Section 4. Gas/Condensate BP Multiphase Design Manual
where:
maximum time interval between spheres (hrs) slug catcher working volume (bbls)
= gas (mmscfd) = liquid loading
The design diagram shows that, although the sphered liquid volume tends towards zero as the gas increases, the sphering operation is still predicted to produce a liquid slug even at high flowrates. However, field data at high flowrates has indicated that above a certain critical gas velocity no slug is produced on sphering. An empirical correlation based on field data gives:
where:
Critical gas velocity (m/s) D = pipe diameter (m)
= gas density = liquid density = surface tension (N/m)
The discrepancy between the present model and field measurements at these high flowrates may be due to the model under-predicting the liquid entrainment in the gas flow (see Section liquid leakage passed the sphere (see Section 4.4) or an underprediction of the liquid film velocity.
4.2.5 Corrosion Inhibition
Corrosion inhibition, using chemicals such as filming is necessary in gas-condensate systems when the required corrosion allowance for an uninhibited system is excessive. If this is the case a corrosion allowance for the inhibited system can be calculated from the allowance for the uninhibited system and the inhibitor efficiency (typically 85 per cent). If the required corrosion allowance is still excessive a higher grade material must used for the pipeline.
Calculation of corrosion allowance and choice of inhibitor is conducted by the Materials Engineering Group.
Corrosion inhibitors will only perform at their nominal if they remain in contact with the pipe wall. However, under adverse flow conditions inhibitor stripping can occur. For single phase flow inhibitor effectiveness is only guaranteed in straight, horizontal flowlines below a crit-ical shear stress value. Materials and Inspection Engineering Group in their Corrosion Guidleline Document [See Section 3, recommend a maximum wall shear rate to avoid inhibitor stripping of 100
The critical shear stress limit is also applied to multiphase systems, where it can be compared with a nominal shear stress value determined from the calculated frictional pressure gradient:
BP Multiphase Design Manual Section 4. Gas/Condensates
D 4 dx
where:
nominal wall shear stress (Pa) D pipe diameter (m)
= frictional pressure gradient (Pa/m) dx
This relationship between shear stress and frictional pressure gradient is based on homoge-neous flow. It always provides a conservative estimate for the actual shear stress in
a steady stratified flow the regime in which gas-condensate flowlines primarily operate.
However, the wall shear stress could be substantially higher when slug flow is encountered, on the outside of bends or on the downstream side of weld beads. To estimate the shear stress-es in thstress-ese locations computational fluid dynamics using the FLUENT code can be employed.
The Multiphase Flow Skills Group have extensive experience in its application.
At present BP has an ongoing programme of research into the factors which dictate corrosion inhibitor performance in multiphase flow. This work may indicate that inhibitor effectiveness is dependent on factors additional to the shear stress.
As gas-condensate pipelines operate primarily in stratified flow regime, corrosion inhibitor is only deposited on the pipe’s top wall if there is a substantial fraction of the liquid entrained as droplets in the the gas phase. However, corrosion will still occur due to the condensation of water. For an uninhibited system the rate of corrosion at the top of the line will be lower than for the of the base of the pipe because, after condensing, the water will rapidly fall back to the stratified liquid layer under gravity. De Waard and Milliams (1) have indicated that a conserva-tive estimate of top of line corrosion is 10 per cent of the overall uninhibited corrosion rate. This has been confirmed by experimental work conducted at
A top of line to bottom of line corrosion rate ratio of 0.1, for an uninhibited system, implies that corrosion considerations will be dictated by the bottom of line rate for inhibited systems, unless the inhibitor efficiency is better than 90 per cent.