Gas Dehydration - Chapter 5 - Part 4

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Gas Dehydration - Chapter 5 - Part 4

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Fundamentals of Oil and Gas Processing Book
Basics of Gas Field Processing Book
Prediction and Inhibition of Gas Hydrates Book
Basics of Corrosion in Oil and Gas Industry Book

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5.22.11.4 Regeneration Gas Separator
Most desiccants also have an affinity for hydrocarbons, thus a skimmer is used to separate the valuable hydrocarbons from the water to be discarded. Frequent pH tests on the discarded water helps pinpoint corrosion problems in the adsorption system. A common problem encountered in regeneration gas separators is the fouling of the liquid dump line by desiccant dust and heavy oils.

5.22.11.5 Expander Plant Molecular Sieve Applications
Turbo expander plants commonly operate down to temperatures of -150 0F.
Operating points much below the equilibrium water content data illustrated in McKetta-Wehe chart (include designs to water contents as low as 1 ppm).
As shown in Table 5-7 (UP), only molecular sieves and activated alumina are capable of such performance. Molecular sieves are used in approximately 95% of the dehydration equipment for this type of plant (a 4A molecular sieve has twice the adsorptive capacity of activated alumina).
5.23 Desiccant Performance
5.23.1 General Conditions
Desiccants decline in adsorptive capacity at different rates under varying operating conditions.
Desiccant aging is a function of many factors, including:
Number of cycles experienced
Exposure to any harmful contaminants present in the inlet stream.
The most important variable affecting the decline rate of desiccant capacity is the chemical composition of the gas or liquid to be dried.

Capacity of a new desiccant will decline slowly during the first few months in service because of cyclic heating, cooling, and netting. Desiccant capacity usually stabilizes at about 55 to 70% of the initial capacity.

5.23.2 Moisture Analyzer
The moisture analyzer is used to optimize the drying cycle time, where the drying time is always shortened as the desiccant ages. Both inlet and outlet moisture analyzer probes should be used. A probe extending approximately 2 feet upward into the bed from the outlet end is recommended because it allows a dehydration capacity test to be run without the risk of a water breakthrough.

5.23.3 Effect of Contaminants in Inlet Feed Stream
Compressor oils, corrosion inhibitors, glycols, amines, and other high-boiling contaminants cause a decline in desiccant capacity, because normal reactivation temperatures will not vaporize the heavy materials. Residual contaminants slowly build up on the desiccant’s surface reducing the area available for adsorption. Many corrosion inhibitors chemically attack certain desiccants, permanently destroying their usefulness.

5.23.4 Effect of Regeneration Gases Rich in Heavy Hydrocarbons
Use of this rich gas in a 5500 to 600 0F regeneration service aggravates coking problems. Rich gases may be dried satisfactorily with molecular sieves. Lean dry gas is always preferable for regeneration.

5.23.5 Effect of Methanol in the Inlet Gas Stream
Methanol in the inlet gas is a major contributor to the coking of molecular sieves where regeneration is carried out at temperatures above 550 0F. Polymerization of methanol during regeneration produces dimethyl ether and other intermediates that will cause coking of the beds.
Conversion to ethylene glycol injection, instead of methanol for hydrate control, will increase sieve life and add at least 10% to sieve capacity by removing the vapor phase methanol from the system.

5.23.6 Useful Life
The most common reasons for replacing a bed are loss of adsorbent capacity and unacceptable pressure drop, which usually occur simultaneously. Values for the loss of capacity with time vary considerably, but common values used for molecular sieves in dehydration service are a 35% capacity loss over a 3 to 5 year period or a 50% loss in approximately 1,600 cycles. Typically, a rapid loss occurs in the beginning and a gradual loss thereafter. The adsorbent decays primarily because of carbon and sulfur fouling and caking caused by instability in the clay binder. These effects occur during bed regeneration.
Increased pressure drop is usually caused by breakdown of adsorbent into finer particles and by caking at the top of the bed because of refluxing. Attrition can occur when the pressure is increased or decreased (more than the allowable pressure-changing rate) after or before regeneration.
Monitoring the pressure drop is important, as it provides a good diagnostic to bed health.
In general, adsorbent life ranges from one to four years in normal service. Longer life is possible if feed gas is kept clean. Effectiveness of regeneration plays a major role in retarding the decline of a desiccant’s adsorptive capacity and prolonging its useful life. If all the water is not removed from the desiccant during each regeneration, its usefulness will sharply decrease.

5.23.7 Effect of Insufficient Reactivation
Insufficient reactivation can occur if the regeneration gas temperature or velocity is too low.
A desiccant manufacturer will generally recommend the optimum regeneration temperature and velocity for the product. Velocity should be high enough to remove the water and other contaminants quickly, thus minimizing the amount or residual water and protect the desiccant.

5.23.8 Effect of High Reactivation Temperature
Higher reactivation temperatures remove volatile contaminants before they form coke on the desiccant, maximizes desiccant capacity and ensures minimum effluent moisture content.
Final effluent hot gas temperature should be held one or two hours to achieve effective desiccant
reactivation.

5.24 Design
5.24.1 Pressure Drop & Minimum Diameter
The first step is to determine the bed diameter, which depends on the superficial velocity. Too large a diameter will require a high regeneration gas rate to prevent channeling. Too small a diameter will cause too high a pressure drop and damage the sieve. The pressure drop is determined by a modified Ergun equation, which relates pressure drop to superficial velocity as follows:
ΔP / L = B ս V + C ρ V2 Eq 5-17

ΔP = pressure drop, psi
L = length of packed bed, ft
B & C = constants from table 5-10
µ= viscosity, cp
V = superficial vapor velocity, ft/min
ρ = density, lb/ft3
Particle Type B C
1/8" bead (4x8 mesh) 0.0560 0.0000889
1/8" extrudate 0.0722 0.000124
1/16" bead (8x12 mesh) 0.152 0.000136
1/16" extrudate 0.238 0.000210
Image
Table. 5-10. Constants for Equation 5-6.

Fig. 5-55 was derived from Eq 5-17 by assuming a gas composition and temperature and setting the maximum allowable ΔP/L equal to 0.33 psi/ft. The design pressure drop through the bed should be about 5 psi. A design pressure drop higher than 8 psi is not recommended as the desiccant is fragile and can be crushed by the total bed weight and pressure drop forces.
Image
Fig. 5-55. Allowable Velocity for Mole Sieve Dehydrator
Remember to check the pressure drop after the bed height has been determined. Once the allowable superficial velocity is estimated, calculate the bed minimum diameter
(i.e., Dminimum), and select the nearest standard diameter (i.e. Dselected):

Dminimum = (4q /π Vmax)0.5 Eq. 5-18

where
D = diameter, ft
q = actual gas flow rate, ft3/min

q = m/60ρ Eq. 5-19

where
m = mass flow rate, lb/hr
ρ= density, lb/ft3

Obtain the corresponding superficial velocity, Vadjusted as follows:
Vadjusted = Vmax (Dminimum/ Dselected)2 Eq. 5-20
An alternative and more exact method to calculate the maximum superficial velocity can be determined by Eq 5-21, which was derived from Eq.5-17.
Vmax = [(ΔP/L)max/(Cρ)]0.5 – [(B/C)(ս/ρ)/2] Eq 5-21

The value of (ΔP/L)max in these equations depends on the sieve type, size, and shape, but a typical value for design is 0.33 psi/ft.

5.24.2 Mass desiccant Required & Bed Length
Choose an adsorption period and calculate the mass of desiccant required. Eight to twelve hour adsorption periods are common. Periods of greater than 12 hours may be justified especially if the feed gas is not water saturated.
Long adsorption periods mean less regenerations and longer sieve life, but larger beds and additional capital investment.
During the adsorption period, the bed can be thought of as operating with three zones.
Saturation or equilibrium zone
The middle or mass transfer zone (MTZ) is where the water content of the gas is reduced from its inlet concentration to < 1 ppm. (lb/MMscf ~ ppmv / 21.4)
The bottom zone is unused desiccant and is often called the active zone. If the bed operates too long in adsorption, the mass transfer zone begins to move out the bottom of the bed causing a “breakthrough.”
In the saturation zone, molecular sieve is expected to hold approximately 13 pounds of water per 100 pounds of sieve. New sieve will have an equilibrium capacity near 20%; 13% represents the approximate capacity of a 3-5 year old sieve.
This capacity needs to be adjusted when the gas is not water saturated or the temperature is above 75°F. Figures. 5-56 and 5.57 are used to find the correction factors for molecular sieve.

Alternatively, both parameters may be calculated using the following corellations:
CSS = 0.636 + 0.0826 ln(Sat) Eq. 5-22
and
CT = 1.20 – 0.0026 t(°F) Eq. 5-23
CT = 1.11 – 0.0047 t(°C) Eq. 5-24

where CSS and CT are correction factors for subsaturation and temperature, respectively.
Sat is the percent of saturation.
To determine the mass of desiccant required in the saturation zone, calculate the amount of water to be removed during the cycle and divide by the effective capacity.

SS = Wr / [(0.13)(CSS)(CT)] Eq 5-25
LS = 4 SS /[π (D2) (bulk density) Eq 5-26
Where
LS = length of packed bed saturation zone, ft
SS = amount molecular sieve required in saturation zone, lb
Wr = water removed per cycle, lb
CSS = saturation correction factor for sieve (Fig.5-56)
CT = temperature correction factor (Fig.5-57)
LS = length of packed bed saturation zone, ft
D = diameter, ft

Molecular sieve bulk density is 42 to 46 lb/ft3 for spherical particles and 40 to 44 lb/ft3 for extruded cylinders.
Even though the MTZ will contain some water (approximately 50% of the equilibrium capacity), the saturation zone is estimated assuming it will contain all the water to be removed.
The length of the mass transfer zone can be estimated as follows:

LMTZ = (Vadjusted/35)0.3 (ZL) Eq 5-27

Where:
LMTZ = length of packed bed mass transfer zone, ft
ZL = 1.70 ft for 1/8 inch sieve
= 0.85 ft for 1/16 inch sieve
Image
Fig. 5-56. Mole Sieve Capacity Correction for gas percent saturation
Image
Fig. 5-57. Mole Sieve Capacity Correction for Temperature

The total bed height is the summation of the saturation zone and the mass transfer zone heights. It should be no less than the vessel inside diameter, or 6 feet, whichever is greater.
Now the total bed pressure drop is checked. The ΔP/L for the selected diameter, Dselected, is adjusted using Eq 5-17 or the following approximation:
(ΔP / L)adjusted ≅ (0.33 psi/ft) (Vadjusted/Vmax)2 Eq 5-28

The result is multiplied times the total bed height (LS +LMTZ) to get the total design pressure drop, which should be 5- 8 psi. This is important, because the operating pressure drop can increase to as much as double the design value over three years. Too high a pressure drop plus the bed weight can crush the sieve. If the design pressure drop exceeds 8 psi, the bed diameter should be increased and the sieve amount and vessel dimensions recalculated.
To estimate the total cylindrical length of a tower, add 3 feet to the bed height, which provides the space for an inlet distributor and for bed support and hold-down balls under and on top of the sieve bed.

5.24.3 Regeneration Calculations
The first step is to calculate the total heat required to desorb the water and heat the desiccant and vessel. A 10% heat loss is assumed.
Qw = (1800 Btu/lb) (lbs of water on bed) Eq 5-29
Qsi = (lbs of sieve)(0.24 Btu/lb 0F) (Trg – Ti) Eq 5-30
Qst = (lbs of steel)(0.12 Btu/lb 0F) (Trg – Ti) Eq 5-31
Qhl =heat loss = (Qw +Qsi+Qst) (0.1) Eq 5-32

where
Qw = desorption of water heat duty, Btu
Qsi = duty required to heat mole sieve to regeneration temperature, Btu
Qst = duty required to heat vessel and piping to regeneration temperature, Btu
Qhl = regeneration heat loss duty, Btu
Trg = regeneration temperature, 0F
Ti = initial, temperature, 0F

The temperature, Trg, is the temperature to which the bed and vessel must be heated based on the vessel being externally insulated (i.e., no internal insulation which is usually the case). This is about 50°F below the temperature of the hot regeneration gas entering the tower.
The weight of the vessel steel is estimated from equations 5-33 and 5-34.
The design pressure, Pdesign, is usually set at 110% of the maximum operating pressure. The value of 0.125 in Eq 5-34 is the corrosion allowance in inches. The term 0.75Dbed is to account for the weight of the tower heads. The value of "3" provides the space for the inlet distributor and support and hold-down balls.
t(inches) = (12DbedPdesign) / (37,600 – 1.2Pdesign) Eq 5-33

Weight of steel (lb) = 155 (t + 0.125) (LS + LMTZ + 0.75Dbed + 3)Dbed Eq 5-34

For determination of the regeneration gas rate, calculate the total regeneration load from Eq. 5-35
Qtr = (2.5)(Qw + Qsi + Qst + Qhl) Eq. 5-35
where
Qtr = total regeneration heat duty, Btu
The 2.5 factor corrects for the change in temperature difference (in – out) across the bed with time during the regeneration period. It assumes that 40% of the heat in the gas transfers to the bed, vessel steel, and heat loss to atmosphere; and the balance leaves with the hot gas.

The regeneration-gas flow rate is calculated from Eq 5-36 below. The heat capacity, Cp, is calculated with Eq 5-37, with the enthalpies obtained from the enthalpy vs. temperature plots for various pressures in (GPSA- Section 24 – or refer to definition of (Cp) in eq. 5-37). The temperature, Thot, is 50°F above the temperature, Trg, to which the bed must be heated. The temperature, Tb, is the bed temperature at the beginning of regeneration, which is the same as the dehydration-plant feed temperature.
The heating time is usually 50% to 60% of the total regeneration time which must include a cooling period. Figure 5-59 shows a typical temperature profile for a regeneration period (heating and cooling). For 8 hour adsorption periods, the regeneration normally consists of 4 1/2 hours of heating, 3 hours of cooling and 1/2 hour for standby and switching. For longer periods the heating time can be lengthened as long as a minimum pressure drop of 0.01 psi/ft is maintained to ensure even flow distribution across the bed.
mrg = Qtr/[Cp(Thot -Tb)(heating time)] Eq 5-36

Cp (Btu/lb/°F) = (Hhot - Hi)/(Thot -Tb) Eq 5-37
Where
mrg = regeneration mass flow rate, lb/hr
Thot = Hot gas temperature, 0F
Tb = bed starting temperature, 0F
Cp = gas heat capacity, Btu/(lb . °F) = (Hhot - Hi)/(Thot -Tb) (Extract the value from fig. 5-60 ,which include values at 600 psia, or (Use Cp = 0.66 as an approximate value for natural gas in regeneration operation ranges) Hhot = enthalpy, BTU/lb, and Hi = enthalpy, BTU/lb. (Other values can be extracted from curves in GPSA, chapter 24 “Total Enthalpy of Paraffin Hydrocarbon Vapor”).
The superficial velocity of the regeneration gas is calculated from Eq 5-38 for which (q) is calculated from Eq 5-19
V = 4q / (π D2) Eq 5-38
where
V = Velocity, ft/sec
D = diameter, ft
q = actual gas flow rate, ft3/min
The calculated superficial velocity can not be less than the value that corresponds with a minimum bed pressure drop of 0.01 psi/ft. This can be determined from Fig. 5-58, which was derived from Eq 5-17 by assuming a gas composition and temperature and setting ΔP/L equal to 0.01 psi / ft.
If the calculated velocity is less than this, the regeneration gas rate, (mrg) must be increased by multiplying it by the ratio Vmin / V, and the period of regeneration should be decreased by multiplying it times the ratio V / Vmin. A more exact method for calculating the minimum superficial velocity is to use Eq 5-21, but to consider it in terms of (ΔP/L)min and Vmin instead of (ΔP/L)max and Vmax.
Image
Fig. 5-58. Minimum Regeneration Velocity for Mole Sieve Dehydrator
Image
Fig. 5-59. Inlet and Outlet Temperatures During Typical Solid Desiccant Bed Regeneration Period
5.24.4 Design Example
100 MMscfd of natural gas with a molecular weight of 18 is water saturated at 600 psia and 100°F and must be dried to –150°F dew point.
Determine the water content of the gas, and the amount of water that must be removed; and do a preliminary design of a molecular-sieve dehydration system consisting of two towers (one dehydration and the other regeneration in a cycle basis). Use 4A molecular sieve of 1/8" beads (i.e., 4x8 mesh). Compressibility factor z = 0.93.
The regeneration gas is part of the plant’s residue gas, which is at 600 psia and 100°F and has a molecular weight of 17. The bed must be heated to 500°F for regeneration. Compressibility factor = 0.95.

Solution Steps
1. Determine the bed diameter and the corresponding ΔP/L and V. First determine the maximum superficial velocity from Eq 5-10. Let the maximum ΔP/L be 0.33 psi/ft.
Vmax = [(ΔP/L)max/(Cρ)]0.5 – [(B/C)(ս/ρ)/2] Eq 5-21

Fro chapter 1, Eq. 1-9
ρg= 0.093 ((MW)P)/TZ lb/ft3 (Eq. 1-19)
ρg = 0.093 x 18 x 600 / (560 x 0.93)
ρ = 1.93 lb/ft3
Or,
ρ = density of gas (ρ)= (18 mole weight) (600 psia) / [10.73 (560 °R)(0.93 z)] = 1.93 lb/ft3
µ = 0.015 centipoise (Fig. 1-11)
C = 0.000089 from table 5-10.
Vmax = {(0.33 psi/ft) / (0.0000889)(1.93) lb/ft3)}0.5 – [(0.056 / 0.0000889) (0.015 centipoise/1.93 lb/ft3)/2]
= 41.4 ft/min
Mass flow rate (m) = (100 X 106 scf/day) (18 lb/lb mole) / [(24 hr/day)(379.5 scf/lb mole)]
m = 198,000 lb/hr of wet gas
From Eq. 5-19
q = (198,000 lb/hr)/((60 min/hr)(1.93 lb/ft3)) = 1710 ft3/min of wet gas
From eq. 5-18
Dminimum = [(4(1710 ft3/min))/(3.14 X 41.4 ft/min)]0.5 = 7.25 ft

Round off upward to 7.5 ft diameter, for which V and ΔP/L are adjusted as follows (eq. 5-20):
Vadjusted = (41.4)(7.25/7.5)2 = 38.7 ft/min
(ΔP/L)adjusted = 0.33(38.7/41.4)2 = 0.29 psi/ft

2. Estimate the amount of water to be removed from the feed per cycle for each bed.
Base this on a 24-hour cycle consisting of 12 hours adsorbing and 12 hours regenerating (heating, cooling, standby, and valve switching).
From chapter 4, fig. 4-8, the water content at 600 psia and 100°F is 88 pounds/MMscf. The water content at a dew point of –150°F is essentially zero, so the water removed is the following:
w = (88-0 lb/MMscf)(100 MMscf/day) / (24 hr/day)
= 367 lb/hr of water removed
Wr = (367 lb/hr) (12 hr) = 4404 lb water removed per 12-hour drying period or 24-hour cycle per bed.

3. Determine the amount of sieve required and the bed height based on a sieve bulk density of 45 lb/ft3 (table 5-7). Since the feed gas is water saturated, the relative humidity is 100%, so CSS is 1.0, and CT is 0.93 at 100°F from Figures. 5-56 & 5-57.
Applying the equations 5-25:
SS = Wr / [(0.13)(CSS)(CT)] Eq 5-25
SS = (4404) / ((0.13)(1.0)(0.93)) = 36,427 lb of sieve for each bed.
From eq. 5-26.
LS = 4 SS /[π (D2) (bulk density) Eq 5-26
LS = (4)(36,427) /((3.1416) (7.5)2 (45)) = 18.3 ft bed height
from eq-5-27
LMTZ = (Vadjusted/35)0.3 (ZL) Eq 5-27
LMTZ = (38.7/35)0.3 (1.7) = 1.8 ft for mass-transfer zone
LS + LMTZ = 18.3 + 1.8 = 20.1 ft of sieve for each bed

The total sieve = (20.1/18.3)(36,427) = 40,010 lb for each bed

4. Check the bed design and pressure drop which is the ΔP/L calculated in Step 1 times the total bed height calculated in Step 3:
(0.29 psi/ft)(20.1 ft) = 5.8 psi which meets the criterion of not exceeding 8 psi.

5. Calculate the total heat required to desorb the water based on heating the bed and vessel to 500°F. First calculate the weight of steel from Eq 5-33 and 5-34. Let the design pressure, Pdesign, be 110% of the operating pressure: Pdesign = (600)(1.1) = 660 psia.
t(inches) = (12DbedPdesign) / (37,600 – 1.2Pdesign) Eq 5-33
Weight of steel (lb) = 155 (t + 0.125) (LS + LMTZ + 0.75Dbed + 3)Dbed Eq 5-34

t = (12)(7.5)(660) / (37,600 – (1.2)(660)) = 1.614 inches
Weight of steel = (155) (1.614 + 0.125) (18.3 + 1.8)+ (0.75) (7.5) + 3) (7.5) = 58,070 pounds
From the following equations:
Qw = (1800 Btu/lb) (lbs of water on bed) Eq 5-29
Qsi = (lbs of sieve)(0.24 Btu/lb 0F) (Trg – Ti) Eq 5-30
Qst = (lbs of steel)(0.12 Btu/lb 0F) (Trg – Ti) Eq 5-31
Qhl =heat loss = (Qw +Qsi+Qst) (0.1) Eq 5-32
Qtr = (2.5)(Qw + Qsi + Qst + Qhl) Eq. 5-35

Qw = (1800 Btu/lb (4404 lb water) = 7,927,000 Btu
Qsi = (40,010 lb) (0.24 Btu/lb/°F) (500°F – 100°F) = 3,841,000 Btu
Qst = (58,070 lb) (0.12 Btu/lb/°F) (500°F – 100°F) = 2,787,000 Btu
Qhl = (2,787,000 + 7,927,000 + 3,841,000) (0.10) = 1,455,000 Btu
Qtr = (2.5) (2,787,000 + 7,927,000 + 3,841,000 + 1,455,000) = 40,025,000 Btu

6. Calculate the flow rate of regeneration gas using Eq 5-36. Let the heating time be 60% of the total regeneration period, and calculate the gas heat capacity, use Cp, = 0.66 (Btu/lb/°F), or (calculate it from Eq 5-37 using enthalpy curves in GPSA, chapter 24 “Total Enthalpy of Paraffin Hydrocarbon Vapor”):
mrg = Qtr/[Cp(Thot -Tb)(heating time)] Eq 5-36
Cp (Btu/lb/°F) = (Hhot - Hi)/(Thot -Tb) Eq 5-37

(60%) (12 hr) = 7.2 hours heating
Cp (@ 600 psia from fig 5-60) = ((545 – 250) (Btu/lb))/((550 – 100) (°F)) = 0.66 Btu/lb/°F

mrg = (40,025,000 Btu)/((0.66 Btu/lb/°F) (550 – 100) (°F) (7.2 hr)) = 18,717 lb/hr (Eq 5-22)

7. Check that the ΔP/L ≥ 0.01 psi/ft at 550°F.
ρg = 0.093 x 17 x 600 / (1110 x 0.95)
ρ = 0.9 lb/ft3

From Eq. 5-19 q = m/60ρ
= 18,717 / (60 x 0.9) = 346.6 ft3/min of hot regeneration gas

Rearranging Eq 5-38:
V = 4q / (π D2) Eq 5-38
V = 4q/πD2 = ((4)(346.6)/((3.414)(7.5)2) = 7.21 ft/min
μ = 0.023 cP (Fig. 1-11)
From ΔP / L = B ս V + C ρ V2 Eq 5-17
ΔP/L = (0.056) (0.023) (7.21) + (0.0000889) (0.9) (7.21)2 = 0.013 psi/ft (Eq 5-6)
This is safely above the minimum value of 0.01 psi/ft needed to prevent channeling.

8. The design results are summarized as follows:
Number of vessels: two
Vessel design pressure and temperature: 660 psig and 600°F
Vessel dimensions: 90 inches (7.5 feet) ID by 23.1 feet tan to tan
Weight of molecular sieve: 2x40,010 lb
Regeneration gas rate: 18,717 lb/hr (10.026 MMscfd)
Regeneration gas temperature: 550°F
Cycle time: 24 hours, 12 hours adsorption, 12 hours regeneration
[img]ttp://oilprocessing.net/data/documents/V5-60.png[/img]
Fig.5-60. Total Enthalpy of Paraffin Hydrocarbon Vapor @ 600 psia.

5.25 Nonregenerable Dehydrator
In some situations, such as remote gas wells, use of a consumable salt desiccant, such as CaCl2, may be economically feasible.

5.25.1 Calcium Chloride Dehydrator Unit
Calcium chloride (CaCI2) dehydrator consists of three sections (Figure 5-61):
Inlet gas scrubber
Brine tray
Solid brine particles

5.25.2 Principles of Operation
Solid desiccant is placed in the top of the unit. Water-wet gas contacting the solid CaCI2 gives up part of its water to form liquid brine to drip down and fill the trays.
Solid anhydrous CaCl2 combines with water to form various CaCl2 hydrates (CaCl2 .XH2O). As water absorption continues, CaCl2 is converted to successively higher states of hydration eventually forming a CaCl2 brine solution.
Inlet gas coming up through the specially designed nozzles on the trays contact the brine efficiently. The wettest gas contacts the most dilute brine (about 1.2 specific gravity).
Approximate 2.5 Lb H2O/lb CaCI2 is removed in the trays. Brine gravity on the top tray is about 1.4. Another 1 lb H2O/lb CaCI2 is removed in the solid bed section.
Outlet water contents of 1 lb/MMscf have been achieved with CaCl2 dehydrators. Typical CaCl2
Maximum dew point depression of 60 0 to 70 0F occurs in this section. Typically used in remote, small gas fields without heat or fuel. capacity is 0.3 lb CaCl2 per lb H20. Superficial bed velocities are 20-30 ft/min and length to diameter ratio for the bed should be at least 3 to 4:1.
CaCl2 dehydrators may offer a viable alternative to glycol units on low rate, remote dry gas wells. The CaCl2 must be changed out periodically. In low capacity – high rate units this may be as often as every 2-3 weeks. Brine disposal raises environmental issues. In addition, under certain conditions the CaCl2 pellets can bond together to form a solid bridge in the fixed bed portion of the tower. This results in gas channeling and poor unit performance.

Advantages Disadvantages
Simple
No moving parts
No heat required
Does not react with H2S or CO2
Can dehydrate hydrocarbon liquids Batch process
Emulsifies with oil
Unreliable
Limited dew point depression
Image
Table. 5-11. Advantages and disadvantage of Calcium Chloride Dehydrator Unit

Operating Problems
Bridging and channeling is a problem.
Brine can crystallize at 85 0F, thus during low flow periods can plug vessel outlet or trays.
Brine carry-over can cause severe corrosion problems.

Image
Fig. 5-61. CaCl2 dehydrator.

Design Considerations
Figure 5-62 illustrates the water content of natural gas dried by solid calcium chloride bed units.
Image
Fig. 5-62 Water content on natural gas driven by CaCl2 unit (Left: freshly recharged; right: just prior to recharging).
5.26 Dehydration by Refrigeration
The dehydration of natural gas can also be achieved by refrigeration and/or cryogenic processing down to – 150°F in the presence of methanol hydrate and freeze protection. The condensed water and methanol streams decanted in the cold process can be regenerated by conventional distillation or by a patented process called IFPEX-1®.
In the latter process illustrated in schematic form in Figure 5-63 a slip stream of water saturated feed gas strips essentially all the methanol in the cold decanted methanol water stream originating in the cold process at feed gas conditions to recirculate the methanol to the cold process. The water stream leaving the stripper contains generally less than 100 ppm wt of methanol. No heat is required for the process and no atmospheric venting takes place.
The process has several major advantages:
• It can obtain dew points in the −100 to −150°F (−70 to –100°C) range.
• It requires no heat input other than to the methanol regenerator.
• It requires no venting of hydrocarbon-containing vapors.
However, it requires external refrigeration to cool the gas, and minimal methanol losses occur in the stripper.

Image
Fig. 5-63. Example IFPEX-1 . Dehydration Process Flow Diagram


5.27 Dehydration by Membrane Permeation
Membranes can be used to separate gas stream components in natural gas such as water, CO2 and hydrocarbons according to their permeabilities. Each gas component entering the separator has a characteristic permeation rate that is a function of its ability to dissolve in and diffuse through the membrane.
The driving force for separation of a gas component in a mixture is the difference between its partial pressure across the membrane. As pressurized feed gas flows into the metal shell of the separator, the fast gas component, such as water and CO2, permeate through the membrane. This permeate is collected at a reduced pressure, while the non-permeate stream, i.e., the dry natural gas, leaves the separator at a slightly lower pressure than the feed.
The amount of methane and other natural gas components in the permeate stream is dependent on pressure drop and the surface area of the membranes. However, 5–10% of the feed stream is a realistic figure. Dehydration by membrane permeation is therefore normally only considered for plants that can make use of low pressure natural gas fuel.

Membranes are characteristic by lightweight, large turndown ratio, and low maintenance, that make them competitive with glycol units in some situations.
Feed pretreatment is a critical component of a membrane process. The inlet gas must be free of solids and droplets larger than 3 microns.
Inlet gas temperature should be at least 20°F (10°C) above the dew point of water to avoid condensation in the membrane.
Units operate at pressures up to 700 to 1,000 psig (50 – 70 barg) with feed gases containing 500 to 2,000 ppmv of water (lb/MMscf ~ ppmv / 21.4). They produce a product gas stream of 20 to 100 ppmv and 700 to 990 psig (48 to 68 barg). The low-pressure (7 to 60 psig [0.5 to 4 barg]) permeate gas volume is about 3 to 5% of the feed gas volume.
This gas must be recompressed or used in a low-pressure system such as fuel gas.
Smith (2004) suggests that membranes used for natural gas dehydration are economically viable only when dehydration is combined with acid-gas removal.
On the basis of commercial units installed and several studies; (Bikin et al., 2003), membranes are economically attractive for dehydration of gas when flow rates are less than 10 MMscfd (0.3 MMSm3/d). Binci et al. claim that membrane units are competitive with TEG dehydrators on offshore platforms at flows below 56 MMscfd (1.6 MMSm3/d). Certainly, the reliability and simplicity of membranes make them attractive for offshore and remote-site applications, provided the low-pressure permeate gas is used effectively. An added benefit compared with TEG units is the absence of BTEX emissions with membranes.
5.28 Other Processes
The first process is the Twister technology, which is discussed in Chapter 6. It has been considered attractive in offshore applications for dehydration because of its simplicity (no moving parts) along with its small size and weight.
Brouwer et al. (2004) discuss the successful implementation on an offshore platform. Some offshore field pressures are greater than 2,000 psi (140 bar), so recompression is not needed with the unit where overall pressure drop is 20 to 30%.
The second process is the vortex tube technology, which also is discussed in Chapter 6. It also has no moving parts. According to vendor information, it is used in Europe in conjunction with TEG addition to remove water from gas stored underground. We found no examples of its use in gas plants.
5.29 Comparison of Dehydration Processes
A number of factors should be considered in the evaluation of a dehydration process or combination of processes. If the gas must be dried for cryogenic liquids recovery, molecular sieve is the only long-term, proven technology available. It has the added advantage that it can remove CO2 at the same time. If CO2 is being simultaneously removed, because water displaces CO2, the bed must be switched before the CO2 breaks through, which is before any water breakthrough.

Enhanced TEG regeneration systems may begin to compete with molecular sieve. Skiff et al. (2003) claim to have obtained less than 0.1 ppmv water by use of TEG with a modified regeneration system that uses about 70% of the energy required for molecular sieves.

High inlet water-vapor concentrations make molecular sieve dehydration expensive because of the energy consumption in regeneration. Two approaches are used to reduce the amount of water going to the molecular sieve bed. First, another dehydration process, (e.g., glycol dehydration) is put in front of the molecular sieve bed. The second option is to have combined beds with silica gel or activated alumina in front of the molecular sieve. The bulk of the water is removed with the first adsorbent, and the molecular sieve removes the remaining water. This configuration reduces the overall energy required for regeneration.

If dehydration is required only to avoid free-water formation or hydrate formation or to meet the pipeline specification of 4 to 7 lb/MMscf (60 to 110 mg/Sm3), any of the above-mentioned processes may be viable. Traditionally, glycol dehydration has been the process of choice.

System constraints dictate which technology is the best to use. Smith (2004) provides an overview of natural gas dehydration technology, with an emphasis on glycol dehydration.
When considering susceptibility to inlet feed contamination, one should keep in mind that replacing a solvent is much easier and cheaper than changing out an adsorbent bed. However, prevention of contamination by use of properly designed inlet scrubbers and coalescing filters, if required, is the best solution.
In a conventional gas plant, where inlet fluctuations are handled in inlet receiving, feed contamination is generally limited to possible carryover from the sweetening unit. However, in field dehydration the possibility exists of produced water, solids, oil, and well-treating chemicals entering the dehydrator.
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Re: Gas Dehydration - Chapter 5 - Part 4

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