Unlike the natural gas compressors which rely upon recycling the after-cooled compressed gas to compressor suction for surge prevention, the surge control in refrigeration systems is quite different because the refrigerant vapors are compressed only to their saturation pressure for dew point temperature, which when cooled for recycling immediately condense into a liquid. Consequently, the hot compressed refrigerant vapors, also known as the hot gas, before any cooling are recycled to compressor suction to make up for any loss of the process chilling load that reduces the amount of available refrigerant vapors to satisfy the minimum flow to the compressor suction, thus avoiding the phenomena known as “surge”.

Three-Stage Refrigeration Cycle

A three-stage propane refrigeration system [1, 2] is adapted into Figure 1 to illustrate a typical refrigeration cycle.

”]As shown in Figure 1, all three process chillers are located on the 1st stage and there is no process side chilling load at the two economizer levels. The shown schematic is a simplest illustration of a three-stage propane refrigeration cycle [1, 2].

The compressed refrigerant vapors are condensed with ambient air. The high pressure (D-0116) and low-pressure (D-0115) economizers separate the flashed vapors that enter the third and second stages of the refrigerant compressor, respectively. To protect the compressor from any liquid carry over, each stage has a suction drum, and all vapors entering the compressor must pass through these drums (D-0112, D-0113 and D-0114).

Surge Protection with Hot Gas and Quench Liquid

So how do you provide a cool vapor stream to prevent the refrigerant compressor from surging in order to keep the discharge temperature from getting too high? To address this challenge, hot compressed vapors from refrigerant compressor discharge are used to vaporize liquid refrigerant, in a similar manner to how the process stream is used to vaporize refrigerant in the chillers.

In the refrigeration cycle of Figure 1, the hot gas from compressor discharge flows into each suction drum as determined by the anti-surge controller to prevent each compression stage from surging. To cool the hot recycled vapors, a temperature controller introduces quench liquid directly from the refrigerant accumulator D‑0111 into the inlet of each compressor suction drum. Ideally, if the compressor anti-surge vapor flow and temperature control valve on quench liquid are perfectly matched and there is no time lag within the control system, the configuration of Figure 1 will work perfectly.

Unfortunately, if there is any mismatch, which in a real world situation always exists,   either too much quench liquid is added to the inlet stream whereby there is liquid level in the suction drums or not enough quench liquid is added whereby the discharge temperature keeps rising. In either mismatch case, the refrigerant compressor will shut down either from high liquid level in the suction drum or from high discharge temperature caused by the high inlet temperature due to recycle of hot gas. 

Surge Protection with Sparged Hot Gas using No Quench Liquid

A well-proven alternative system is shown in Figure 2 [3, 4] in which hot propane vapors from the compressor discharge, as controlled by the anti-surge system, flow through the sparged line submerged in the liquid filled portions of the HP (D-0116) and LP (D-0115) economizers, and under the tube bundle of the first stage chillers (E-0107, E-0108, and E-0111). There is no need for quench liquid as the hot vapors provide the heat to the available liquid propane in these services to generate the required propane vapors to prevent the compressor surge conditions.

”]The configuration of Figure 2 builds upon the recognition that:

• The shortage of propane refrigerant entering the compressor stages is caused by reduction in heat exchanger duty from changes on the process side of the chillers (E-0107, E-0108 or E-0111). This reduction of heat transfer duty is offset by routing the required amount of hot propane vapor from compressor discharge directly to the kettle side of the chiller to supplement process thermal requirements.
• The compressor suction drums are provided to prevent any carried-over liquid from entering the compressor and should essentially have no liquid in them.
• By diverting the anti-surge hot vapor away from the compressor suction drums assures that the compressor suction drums stay dry.
• By using the already present liquid refrigerant in the chillers and economizers it avoids any potential mismatch between the hot propane vapors and the quench liquid that may inadvertently buildup high liquid level and end up tripping the refrigerant compressor.

Conclusion

While both refrigerant compressor surge control systems are proven and used by the industry, the hot sparged gas with no quench liquid configuration of Figure 2 is more forgiving and reliable.

To learn more, we suggest attending our G40 (Process/Facility Fundamentals), G4 (Gas Conditioning and Processing), G5 (Gas Conditioning and Processing-Special), and PF81 (CO₂ Surface Facilities), PF4 (Oil Production and Processing Facilities), courses.

            John M. Campbell Consulting (JMCC) offers consulting expertise on this subject and many others. For more information about the services JMCC provides, visit our website at www.jmcampbellconsulting.com, or email us at consulting@jmcampbell.com. 

 By: Yuv R. Mehra

References:

1. Al-Shahrani, Saleh M. and Mehra, Yuv R., “Start-up Experience for Value Recovery from CCR Net Gas at Yanbu’ Refinery,” Advances in Hydroprocessing II Session, Paper 26d, 2007 Spring National Meeting of AIChE, Houston, Texas, USA, April 22-26, 2007.
2. Al-Shahrani, Saleh M. and Mehra, Yuv R., “Saudi Aramco installs new LPG recovery unit at Yanbu’ Refinery,” Oil & Gas Journal, Vol. 105.21, Jun. 4, 2007, p. 60.
3. Mehra, Yuv R., “Focus on Value-Chain Contributions during Process Technology Selection,” Worldwide Developments Forum, 86^th Annual Meeting of the Gas Processors Association, San Antonio, Texas, USA, March 11-14, 2007.
4. Mehra, Yuv R., “Focus on Value-Chain Contributions during Process Technology Selection,” Saudi Aramco Journal of Technology, Summer 2007, p. 2.

The details of a simple single-stage refrigeration system, a two-stage refrigeration system employing one flash tank economizer, and with heat exchanger economizer system are given in Chapter 15 of Gas Conditioning and Processing, Volume 2 [1].

In the January 2008 Tip of the Month (TOTM) [2], we compared the performance of a simple refrigeration system with another employing a flash economizer. Specifically, we evaluated compressor power saving, the effects of compressor suction–line pressure drop and the interstage pressure drop on compressor power requirement and condenser duty.

A second type of economizer configuration is the heat exchanger economizer shown in Figure 1, which is the same as Figure 15.9 of reference [1]. Cold, low-pressure chiller vapor is used to subcool the saturated liquid refrigerant. This decreases the refrigerant circulation rate, and may reduce compressor power. In this TOTM we will evaluate quantitatively the performance of a case study comparing a simple refrigeration system with another one containing heat exchanger economizer.

The process flow diagrams for the simple and with heat exchanger (HEX) economizer refrigeration systems are shown in Figure 2. Note that provisions have been made to consider pressure drop in different segments of the loops.

”]Let’s consider removing 1.0×10⁷kJ/h which is equal to 2778 kW (9.479 MMbtu/hr) from a process gas at -35°C (-31°F) and rejecting it to the environment by the condenser at a condensing  temperature of 35°C (95°F). Assuming 5 kPa (0.7 psi) pressure drop in the chiller and 5 kPa in the suction line pressure drop, the compressor suction pressure is 132.4 kPa (19.1 psi). The condenser pressure drop plus the pressure drop in the line from the compressor discharge to the condenser was assumed to be 50 kPa (7.3 psi); therefore, compressor discharge pressure is 1270 kPaa (184.2 psia). The compressor discharge temperature is 66°C (150.8°F). At these conditions, the condenser duty is 4434 kW (15.13 MMbtu/hr). Pure propane is used as the working fluid. In this study, all of the simulations were performed by UNISIM software [3].

Figure 2. Process flow diagrams for a simple refrigeration system and with a heat exchanger economizer

In order to study the effect of HEX economizer, we considered the following scenarios:

1. The condensed liquid at temperature of 35°C (95°F) was cooled starting from 33 to 24 °C (91.4 to 75.2°F) with a step change of -1°C (-1.8°F).
2. Step 1 was repeated three times for pressure drops of  20, 25, and  30 kPa ( 2.9, 3.63,  or 4.4 psi) on both sides of HEX economizer.
3. For the above four cases the following variables were calculated:
4. The required compressor power
5. Compressor suction temperature
6. Compressor discharge temperature
7. Refrigerant (propane) circulation rate
8. Condenser heat duty
9. Chiller inlet temperature

Figures 3, 4, and 5 present the required compressor power, condenser duty, and HEX duty as a function of liquid propane subcooled temperature (at the outlet of HEX economizer), respectively. Figures 3 and 4 indicate that as the propane subcooled temperature decreases the compressor power and condenser duty decrease, too. However as the pressure drop in the cold vapor (low pressure) side increases, the compressor power and condenser duty increase. The pressure drop significantly increases the compressor power. Figure 5 indicates as the propane subcooled temperature decreases, the HEX duty increases independent of  HEX pressure drop.

Figure 3. Compressor power as a function of refrigerant subcooled temperature and HEX economizer pressure drop

Figure 4. Condenser duty as a function of refrigerant subcooled temperature and HEX economizer pressure drop

Figure 5. HEX duty as a function of refrigerant subcooled temperature and HEX economizer pressure drop

The refrigerant mass circulation rate, compressor suction temperature, and compressor discharge temperature as a function of propane subcooled temperature and HEX pressure drop are presented in Figures 6, 7, and 8, respectively. Figure 6 indicates as the propane subcooled temperature decreases, the mass circulation rate decreases independent of HEX pressure drop. Figures 7 and 8 indicate that compressor suction and discharge temperatures increase with decrease in propane subcooled temperature. However, the effect of HEX pressure drop on discharge temperature is more pronounced.

Contrary to a refrigeration system with a flash thank economizer in which the compressor power is reduced [2], by employing HEX economizer the power requirement increased. With regard to the compressor power, two factors offset the reduced circulation rate. The first is HEX pressure drop. The pressure drop on the low pressure side significantly increased compressor power, since the suction pressure was near atmospheric. Secondly, the refrigerant vapor entering the compressor is now super-heated. Although this reduces the likelihood of liquid carryover into the compressor, it resulted in higher power consumption due to the higher suction temperature.

To learn more, we suggest attending our G40 (Process/Facility Fundamentals), G4 (Gas Conditioning and Processing), G5 (Gas Conditioning and Processing-Special), and PF81 (CO₂ Surface Facilities), PF4 (Oil Production and Processing Facilities), courses.

John M. Campbell Consulting (JMCC) offers consulting expertise on this subject and many others. For more information about the services JMCC provides, visit our website at www.jmcampbellconsulting.com, or email us at consulting@jmcampbell.com.

By: Dr. Mahmood Moshfeghian

Figure 6. Refrigerant mass circulation rate as a function of refrigerant subcooled temperature and HEX economizer pressure drop

Figure 7. Compressor suction temperature as a function of refrigerant subcooled temperature and HEX economizer pressure drop

Figure 8. Compressor discharge temperature as a function of refrigerant subcooled temperature and HEX economizer pressure drop

Reference:

1. Campbell, J.M., “Gas conditioning and Processing, Vol 2: The Equipment Modules”, 9^th Edition, Edited by R.A. Hubbard K.S. McGregor, John M. Campbell & Company, Norman, USA, 2014.
2. Moshfeghian, M., “Refrigeration with Flash Economizer vs Simple Refrigeration System,” http://www.jmcampbell.com/tip-of-the-month/2008/01/refrigeration-with-flash-economizer-vs-simple-refrigeration-system/ , 2008
3. UniSim Design, Version 410 Build 17061, Honeywell International, Inc., Calgary, Canada, 2013.

This Tip of the Month describes simple equations to approximate changes to the properties of crude oil with changing temperature.   Changes in crude oil density and specific heat, or heat capacity, can be estimated from graphs and/or more elaborate computer simulation.  The latter generally requires access to a process simulator and characterization data for the crude oil.  A suitable, tuned computer model is likely the most accurate method of estimating the fluid properties, but is not always available.  Direct laboratory measurement is also possible if facilities and oil samples are available and a high degree of accuracy is required.

Graphs, which were originally generated from empirical data, can be useful, and their accuracy is suitable for most engineering applications. However; use of the data in subsequent calculations requires users to interrupt the calculation, look up a number from a graph, and then proceed with the calculation.  The simple curve-fit equations presented provide the required data suitable for use in spreadsheet or hand calculations.

Crude Oil Density

Figure 1A (1) depicts the change in specific gravity with temperature for crude oils of varying API gravity.

Figure 1A Crude Oil Specific Gravity vs Temperature (1)

This graph was compared with density-temperature data from Table D-1, API Publication 421, “Monographs on Refinery Environmental Control – Management of Water Discharges” (2).  The colored lines in Figure 1B show the API data superimposed on the original graph. The two show good agreement.

Figure 1B Data from API Publication 421 superimposed on Figure 1A (data from (2))

Curve fitting data from Figure 1A resulted in Equation 1 for FPS units.  Equation 2 provides the SI equivalent.

Figure 1C displays the output from Equation 1 superimposed as colored lines on the original graph (Figure 1A). Although the simple equation does not align perfectly, results are sufficiently accurate for most engineering calculations.  Compared to the data from API Publication 421,  Equation 1 produces a maximum errors of +0.25% and -0.3%.

Figure 1C Simple Equation 1 superimposed on Figure 1A

Heat Capacity

A similar approach was used to develop a simple equation for the variation of the Heat Capacity or Specific Heat of crude oil as a function of API gravity and Temperature.  Data, extracted from Figure 2A (1), was regressed to obtain the algorithm presented as Equation 3 for FPS units: Equation 4 for SI units.  Note that the algorithm was developed for a crude oil with a UOP index of 11.8 (indicating intermediate, paraffinic-naphthenic crude).  If the UOP index is known, the correction factor illustrated on the graph could be applied to the output from Equation 3 or 4.

Figure 2A Heat capacity of crude oil vs Temperature (1)

The resulting equation is presented as Equation 3 for FPS units, and Equation 4 for SI units.

Figure 2B Simple Equation 3 superimposed on Figure 2A

Summary

The simple equations provide approximations for the variation of density and specific heat of crude oils of varying API gravity.  Neither algorithm provides a perfect match with the underlying data.  However; data from varying sources do not always correlate.  Figure 3A (1) depicts an alternate source of density correction for crude oils for varying API gravity and temperature.  Figure 3B shows data from API Publication 421 (colored lines) superimposed on a portion of Figure 3A.  Unfortunately, overlap of the data is limited, but clearly there is a very poor correlation – even the trend as API gravity increases is reversed between the API data and the data presented in Figure 3A.  Note that the trend represented in Figure 1A is supported by data from API 421 Appendix D.  These data (from the API Publication 421) are taken here as being most reliable.  Overall, Figure 1A shows quite good agreement with the API data, so the algorithm (Equations 1 and 2) was developed using data from Figure 1A as the source.

Figure 3A Crude Oil Density Correction Factor (Hankinson et al, 1979) (1)

Figure 3B Portion of Figure 3 A with data from API 421 superimposed

To learn more about similar cases and how to minimize operational problems, we suggest attending our G40 (Process/Facility Fundamentals), G4 (Gas Conditioning and Processing), PF81 (CO₂ Surface Facilities), and PF4 (Oil Production and Processing Facilities) courses.

John M. Campbell Consulting (JMCC) offers consulting expertise on this subject and many others. For more information about the services JMCC provides, visit our website at www.jmcampbellconsulting.com, or email us at consulting@jmcampbell.com.

By Wes Wright

Works Cited

1. Manning, Francis S. and Thompson, Richard E. OILFIELD PROCESSING Volume Two: Crude Oil. Tulsa : PennWell Publishing Company, 1995. ISBN 0-87814-354-8.

2. Institute, American Petroleum. Management of Water Discharges: Design and Operation of Oil-Water Separtors. Washington : API, 1990. API Publication 421.

In the January, February, and March 2012 tips of the month (TOTM) we discussed the transportation of carbon dioxide (CO₂) in the dense phase region. We illustrated how thermophysical properties changed in the dense phase and studied their impacts on pressure drop calculations. We showed that the effect of the numerical range of values for the overall heat transfer coefficient on the pipeline temperature is significant.

In this TOTM, we will study the transportation of ethane by pipeline in the dense phase region. For a case study, a mixture of ethane containing a small fraction of methane and propane was considered. The pressure and temperature profiles along the pipelines are calculated and plotted on the feed phase envelope. In addition, the pump power requirement, pressure and temperature profiles for a single pipeline with a lead pump station are compared with the option of dividing the same line into three equal segments having one lead pump station and two intermediate pump stations.

Calculation Procedure:

For a pure compound above critical pressure and critical temperature, the system is often referred to as a “dense fluid” or “super critical fluid” to distinguish it from normal vapor and liquid (see Figure 1 for carbon dioxide in December 2009 TOTM [1]).

The same step-by-step calculation procedure described in the February 2012 TOTM [2] was used to determine the pressure and temperature profiles in a pipeline.

In the following section we will illustrate the pressure drop calculations for transporting ethane mixture in the dense phase. For details of pressure drop equations in the gas and liquid phases refer to the January 2012 TOTM [3].

Case Study:

For the purpose of illustration, we considered a case study for transporting 31393 kg/h (69209.0 lb_m/hr) equivalent to 275000 tonne/y of the cited  ethane  mixture with the composition presented in Table 1. The mixture is available at the pressure and temperature presented in Table 1.

The following assumptions were made:

1. Horizontal pipeline, no elevation change
2. Steady state conditions
3. The pipeline is 1380 km (858 mile) long with an inside diameter of 208.6 mm (8.212 in), onshore buried line.
4. Pipeline inside surface roughness of 46 microns (0.046 mm, 0.0018 inch) which is equivalent to inside surface relative roughness (roughness factor), ε/D, of 0.00022.
5. Delivery Pressure is 5.3 MPa  (769 psia)
6. The ground/ambient temperature, is 18.3 ˚C, (65 ˚F)
7. Overall heat transfer coefficient of 2.839 W/m²-˚C (0.5 Btu/hr-ft²-˚F), for onshore buried line.
8. Pump efficiency is 50%, this is the worst case, the actual pump efficiency is in the range of 50-85%).
9. Simulation software: ProMax [4] and using Soave-Redlich-Kwong (SRK) [5] Equation of State.

The block diagrams for two options studied are presented in Figure 1. In option A, only one pump station and a single long segment were considered. In option B, for the same pipeline and feed conditions, one lead and two intermediate pump stations with three equal pipeline segment were considered.  Each segment length is 460 km (286 mile), which is 1/3 of total length and all have the same inside diameter of 208.6 mm (8.212 in).  The delivery pressures for both options are the same.

Results and Discussions:

Figure 2 presents the phase envelope for the ethane mixture with the composition presented in Table 1. According to Figure 2, the feed at the pump suction conditions of 0°C (32°F) and 3000 kPa (435 psia) presented in Table 1 is in the liquid phase.  In order to deliver the ethane mixture at pressure of 5.3 MPa (769 psia), for option A this liquid is pumped to a pressure of 13.6 MPa (1972 psia) before entering the pipeline. Due to pumping, the feed temperature rises from 13.6 MPa (1972 psia) before entering the pipeline. A step-by-step pump calculation with increments of 2 MPa (290 psia) for the discharge pressure reveals that the temperature rise is linear with pressure. This significant temperature rise is due to compressibility of ethane mixture.

Figure 2. Phase envelope for ethane mixture.

The pumping and pipeline pressure-temperature paths for option A are plotted on the phase envelope and presented in Figure 3. As shown in this figure, the ethane mixture at the pump discharge (pipeline inlet) is in the supercritical region (dense phase).  Figure 3 indicates that as the feed enters the pipeline, its temperature drops rapidly and remains constant and very close to the ambient temperature.

Figure 3. Phase envelope, pumping path and dense phase pipeline pressure-temperature profile.

Figure 4 presents the calculated pipeline pressure profiles for options A and B. For option A, the inlet pressure is 13.6 MPa (1972 psia) to assure ethane mixture delivery at 5.3 MPa (769 psia). Similarly, in option B at each pump station the pressure is increased to 8.2 MPa (1189 psia).

Figure 5 presents the calculated pipeline temperature profiles for options A and B. The constant ambient temperature of 18.3˚C and (65˚F) is also plotted. In option B, the discharge temperature for lead pump station is 11.2˚C, (52.1˚F) which is below the ambient temperature. For both options, the pipeline temperature rapidly approaches the ambient temperature within the first 50 km (31 mile).

Figures 6 and 7 present the density and velocity profiles along the pipeline, respectively. For a crude oil cross-country pipeline, the velocity is in the range 1.5 to 2.5 m/s (5 to 8 ft/sec). The abrupt change of density, and consequently velocity, along the first 60 km (37.3 mile) is due to the ethane mixture temperature drop, which approaches the ground temperature.

Figure 4. Pipeline pressure profiles for options A and B.

Figure 5. Pipeline temperature profiles for options A and B.

Figure 6. Pipeline fluid density profile for option A

Figure 7. Pipeline fluid velocity profile for option A

The total pump power requirements with pump efficiency of 50% for options A and B are 457 and 504 kW (381.6 and 420.8 hp), respectively. This is for screening estimate only. Normally, work is done in collaboration with the pump manufacturers for better efficiencies based on CAPEX and OPEX.

Conclusions:

Based on the results obtained for the case study considered in this TOTM, the following conclusions can be made:

1. During the pumping of ethane mixture, the temperature rises linearly with pressure (Figure 2).
2. The feed temperature approaches the ground temperature rapidly (Figure 5). This may not be the case for lower overall heat transfer coefficient.
3.  A single pipeline with only one lead pump station (option A) requires smaller pump power compared to the option of one lead pump and two intermediate pump stations (option B). Due to higher pressure in option A, wall thickness will be higher.
4. A complete cost analysis of pumping requirement vs pipeline cost should be made to determine the optimum pipeline diameter, wall thickness and power requirement.
5. The point not considered but worth mentioning is that ethane is very difficult to seal. We would work with pump and seal manufacturers for selecting the correct dry gas seal. This selection could determine the overall system reliability.

To learn more about similar cases and how to minimize operational problems, we suggest attending our G40 (Process/Facility Fundamentals), G4 (Gas Conditioning and Processing), PF81 (CO₂ Surface Facilities), PF4 (Oil Production and Processing Facilities), and PL4 (Fundamentals of Onshore and Offshore Pipeline Systems) courses.

John M. Campbell Consulting (JMCC) offers consulting expertise on this subject and many others. For more information about the services JMCC provides, visit our website at www.jmcampbellconsulting.com, or email us at consulting@jmcampbell.com.

By: Dr. Mahmood Moshfeghian

Reference:

1. Bothamley, M.E. and Moshfeghian, M., “Variation of properties in the dese phase region; Part 1 – Pure compounds,” TOTM, http://www.jmcampbell.com/tip-of-the-month/2009/12/variation-of-properties-in-the-dense-phase-region-part-1-pure-compounds/, Dec 2009.
2. Moshfeghian, M., ”Transportation of CO₂ in the Dense Phase,” TOTM, http://www.jmcampbell.com/tip-of-the-month/2012/02/ , Feb 2012
3. Moshfeghian, M., ”Transportation of CO₂ in the Dense Phase,” TOTM, http://www.jmcampbell.com/tip-of-the-month/2012/01/, Jan 2012
4. ProMax 3.2, Bryan Research and Engineering, Inc, Bryan, Texas, 2014.
5. Soave, G., Chem. Eng. Sci. 27, 1197-1203, 1972.

In the past Tips of the Month (TOTM), we discussed the phase behavior of sweet natural gas- water, sour natural gas-water, and acid gas–water systems. They were posted in October 2007 TOTM [1], November 2007 TOTM [2], and December 2007 TOTM [3], respectively. In this TOTM, we will revisit the acid gas-water phase behavior system. Specifically, different methods of predicting water content of acid gas systems are evaluated based on experimental data from the literature. Water content diagrams compatible with the experimental data for pure CO₂, Pure H₂S, pure CH₄ and their mixtures are generated and presented. These charts can be used for facility type calculations and trouble shooting.

Table 1 presents the compositions of several acid gas mixtures evaluated in this study, along with their saturated water contents (in mole percent) from experimental data [4] and from predictions by five methods. The Maddox et al. [5] results were generated using GCAP [6] software and the Erbar et al. [7] results were generated by EzThermo [8] software. The Wichert & Wichert [9] and Yarrison et al. [10] results are from GPA RR-210 [11]. The last column presents the results predicted by SRK in ProMax [12].

Table 1 indicates that as long as the total acid gas concentrations is less than 60 mole percent, all five methods produce results within the accuracy of experimental data. However, for higher concentrations of acid gases, the Yarrison et al. [10] and ProMax [12] methods provide more accurate results.

It should be noted that the results of ProMax in Table 1 are based on the water saturator tool available in ProMax. If a conventional 3-phase separator calculation of ProMax is used, the error percent for the third row from the bottom of table reduces from -44.9 % to -16.6 %. The results for the other cases remained practically the same for these two calculations options.

Table 2 and Figure 1 present the error analysis for prediction of 74 additional gas mixtures containing acid gases. The detail of data points and sources of experimental points are in GPA RR-210 [11]. The five methods under study are Maddox et al. [5], Robinson et al. [13], Wichert & Wichert [9], Yarrison et al. [10], and ProMax [12]. With the exception of the ProMax results, the predicted water contents used for error analysis of the other four methods were extracted from   GPA RR-210 [11].

Figure 1 presents the error analysis graphically for the same 74 gas mixtures presented in Table 2. Based on the error analysis of Tables 1, 2, and Figure 1, the ProMax method was chosen to generate water content diagrams for pure CO₂, pure H₂S, and their mixtures. These diagrams are presented in the proceeding sections.

The phase equilibria in the system H₂S + water and CO₂ + water are key to the discussion of the water content of an acid gas system. Figures 2 (SI) and 2 (FPS) present the water content of pure H₂S predicted by ProMax [12] as a function of pressure and temperature, in the international system (SI) and engineering system of units (FPS, foot, pound and second). The behavior shown on this plot is quite complicated and explained thoroughly by Carroll [14]:

“At low pressure the hydrogen sulfide + water mixture is in the gas phase. At low pressure the water content tends to decrease with increasing pressure, which is as expected. Eventually a pressure is reached where the H₂S is liquefied. On this plot this is represented by the discontinuity in the curve and a broken line joins the phase transition. There is a step change in the water content when there is a transition from vapor to liquid. In the case of hydrogen sulfide the water content of the H₂S liquid is greater than the coexisting vapor. This is contrary to the behavior for light hydrocarbons where the water content in the hydrocarbon liquid is less than the coexisting vapor.”

Within the transition region, the acid gas exists as both liquid and vapor.  Water saturation of the vapor phase is represented by the lower value, whereas the water content of the liquid phase is the higher value.

Figures 3 (SI) and 3 (FPS) present the water content of pure CO₂ and pure CH₄ predicted by ProMax [12] as a function of pressure and temperature, in the international system (SI) and engineering system of units (FPS). When Figures 2 and 3 were superimposed on Figures 20-5 and 20-6, respectively, of the GPSA data book [15] a very good match was obtained. The two figures in the GPSA data book are based on experimental data and the Yarrison et al. model.

In general the phase behavior of the system CO₂ + water is as complex as that of the H₂S + water system. The CO₂-rich liquid phase only occurs for temperatures less than about 32.2°C (90°F). As shown in Figure 3 (as well as in Figure 2 reported by Maddox and Lilly [16]), the water content of CO₂ exhibits a minimum.

Figure 4 presents the phase behavior of pure CO₂, Pure H₂S and three mixtures of them containing 2 mole percent CH₄. Their corresponding water content charts are presented in Figure 5.

Summary:

There are several methods available that can be used to predict the water content of acid gases. Most of these methods are based on equations of state and rigorous thermodynamic models. As described above, the phase behavior is complicated and extra care should be taken to assure a correct prediction.  Although not addressed in this study, hydrates can also form and these can significantly complicate phase behavior.

Different methods of predicting water content of acid gas systems are evaluated based on the literature experimental data. In addition, the water content diagrams compatible with the experimental data for pure CO₂, H₂S, CH₄ and their mixture are generated and presented. These charts can be used for facility type calculations and trouble shooting.

To learn more, we suggest attending our G40 (Process/Facility Fundamentals), G4 (Gas Conditioning and Processing), G5 (Gas Conditioning and Processing-Special), and PF81 (CO₂ Surface Facilities), PF4 (Oil Production and Processing Facilities), courses.

John M. Campbell Consulting (JMCC) offers consulting expertise on this subject and many others. For more information about the services JMCC provides, visit our website at www.jmcampbellconsulting.com, or email us at consulting@jmcampbell.com.

References:

1. Moshfeghian, M. “Water-Sweet Natural Gas Phase behavior,” http://www.jmcampbell.com/tip-of-the-month/2007/10/water-sweet-natural-gas-phase-behavior/, October 2007.
2. Moshfeghian, M., ”Water-Sour Natural Gas Phase Behavior,” http://www.jmcampbell.com/tip-of-the-month/2007/11/water-sour-natural-gas-phase-behavior/, November 2007.
3. Wright, W. and M. Moshfeghian, “Acid Gas-Water Phase Behavior,” http://www.jmcampbell.com/tip-of-the-month/2007/12/acid-gas-water-phase-behavior/, December, 2007.
4. Huang, S.S.S., A.D. Leu, H.J. Ng, and D.B. Robinson, “The Phase Behavior of Two Mixtures of Methane, Carbon Dioxide, Hydrogen Sulfide, and Water” Fluid Phase Equil. 19, 21-32, 1985.
5. Maddox, R.N., L.L. Lilly, M. Moshfeghian, and E. Elizondo, “Estimating Water Content of Sour Natural Gas Mixtures”, Laurance Reid Gas Conditioning Conference, Norman, OK, Mar., 1988.
6. GCAP 8.3, John M. Campbell & Co., Norman, Oklahoma, December 2010.
7. Erbar, J.H., A.K. Jagota, S. Muthswamy, and M. Moshfeghian, “Predicting Synthetic Gas and Natural Gas Thermodynamic Properties Using a Modified Soave-Redlich-Kwong Equation of State,” Gas Processor Research Report, GPA RR-42, Tulsa, USA, 1980.
8. EzThermo, Chemical Engineering Consultants, Inc, Stillwater, Oklahoma, 2010.
9. Wichert, G. C. and E. Wichert, “New Charts Provide Accurate Estimation for Water Content of Sour Natural Gas”, O&G J, pp 64-66, Oct. 27, 2003..
10. Yarrison M., Song, K. Y., Cox, K,, Chronister D. and Chapman, W., “Water Content of High Pressure, High Temperature Methane, Ethane and Methane+CO2, Ethane + CO2,” RR-200, GPA, Tulsa, OK, March, 2008.
11. Song, K. Y., Vo, T., Yarrison M. and Chapman, W., “Acid Gas Water Content, An Update Of Engineering Data Book I,” RR-210, GPA, Tulsa, OK, June, 2012.
12. ProMax 3.2, Bryan Research and Engineering, Inc., Bryan, Texas, U.S.A., 2013.
13. Robinson, J. N., et al., Trans. AIME, Vol. 263, p. 281, 1977
14. Carroll, J.J., “The water content of acid gas and sour gas from 100 to 220 °F and pressures to 10,000 Psia,” Presented at the 81st Annual GPA Convention, Dallas, Texas, USA, March 11-13, 2002.
15. GPSA Data Book, Vol. 2, 13^th Ed., Gas Processors and Suppliers Association, Tulsa, Oklahoma, 2013.
16. Maddox, R.N., L.L. Lilly, “Gas conditioning and Processing, Vol 3: Computer Applications and Production/Processing Facilities”, John M. Campbell & Company, Norman, USA, 1982.

A pillar of Risk Based Process Safety (RBPS) is Learn from Experience.  The work we do and the processes we use to analyze our work provide significant learning opportunities to enhance process safety competency.  This is a derivative of Kolb’s experiential learning cycle [1], but many times we fail to take advantage of the learning opportunities available to us unless there is an incident or a near miss.

This Tip of the Month (TOTM) will introduce a simple method for debriefing the job tasks we perform to close the loop on this cycle and capture appropriate data to develop competency, work safely and capture near miss/incident data quickly and efficiently.

Conducting a simplified job hazard analysis will ensure that all hazards are identified, managed, and mitigated prior to performing work.  Performing a simple debrief at the conclusion of the work will ensure that we learn from the experience. By considering every job to be performed a learning opportunity, the experiential learning cycle can be used to identify what was done, how well it was done, and how we might improve in the future.

This Month’s Tip was recently presented at the Mary K. O’Connor Process Safety Symposium at Texas A&M University [1].

One of the pillars of the Center for Chemical Process Safety’s (CCPS) Guidelines for Risk Based Process Safety is “Learn from Experience.”  What does this mean?

The elements of this pillar include:

• auditing,
• management review and continuous improvement,
• measurement and metrics and
• incident investigation.

Each of these elements provides findings, lessons and data that are useful for learning and thus changing and enhancing behaviors and attitudes.  The change and enhancement will influence an organization’s culture and ultimately push the organization toward a learning culture.

These are not the only opportunities available for organizations to learn from experience.  Metrics and audits can allow a general overview of process safety performance.  Incident investigation insures that when reported, incident information is transmitted to all who will benefit from the learning.

The job hazard analysis process that many organizations use to identify and mitigate hazards provides a tremendous opportunity to capture data and use the experiential learning cycle if the job is debriefed properly after completion.  This paper will provide guidance and explain the benefits that can be derived from debriefing completed jobs.

At the 2008 symposium, this author presented a paper entitled “Three Simple Things to Improve Process Safety Management.”  One of those simple things was to conduct a formalized Job Hazard Analysis (JHA) for the tasks being performed in the life cycle of a process.  That paper presented a checklist that could be used to guide personnel in the process of conducting a JHA.  (See checklist at end of this paper)

Many facilities have embraced the concept of conducting JHA.  They may be called something else  (job safety analysis, job safety checklist, job task analysis) but the process is essentially the same.  The job or task is identified and analyzed step by step.  The analysis is to identify hazards that may be involved with each step and then develop strategies to mitigate the hazards.  This sounds simple in theory, but in reality there are many things that can and do go wrong with this process.

To provide consistency and to make it easier to track that these analyses have been completed, standardized checklists and forms have been created that list the most common hazards that can be found with a job and logically guide the user toward identification and mitigation of hazards.  Experience shows that after these forms and checklists have been used regularly, some personnel have a tendency to try and short cut the process.  This leads to what is known as “pencil whipping” the JHA.  In other words, personnel will complete the checklist or form without actually performing the analysis required.   Familiarity with the forms and checklists may drive personnel to identify common hazards, but do little to mitigate the hazards.  For example, a common checklist item is “slips, trips and fall hazards”.  Personnel will identify that the ground is rutted or that there is ice on the ground, but few will actually smooth the ground or cover the ice with sand to mitigate the hazards identified.

It is generally agreed among those who supervise personnel performing JHAs that the most important part of the process is not the completion of the forms and checklists but the discussion that happens among a group performing the work.  In order to focus the discussion and insure that all issues are addressed, the JHA checklist at the end of this paper can be used.  The JHA checklist is not intended to replace the checklists and forms that an organization may already have in place.  The JHA checklist can enhance the process by focusing a group’s thoughts on individual checklist items.  By answering each question a work group should be able to identify all issues associated with any job they are conducting.

As work groups become more familiar with the JHA checklist and the process of discussing and documenting the efforts of the group, a simplified method can be adopted.  By answering six key questions, a group of workers can focus discussion on the issues that are most important.   The six questions and the benefits of using them include:

What are we doing?  If we can’t answer this question completely and in simple terms, then we should not be doing the job.  A simple explanation will insure that all members of the team are working toward the same goal.

What is the most dangerous part?  If we can identify the most dangerous part of what we are doing we have identified all hazards, ranked them and determined the most dangerous part.

What will we do to protect ourselves?  Answering this question ensures that all mitigation measures have been put into place and that all personnel know what is being done.

How will we know we are changing what we are doing?  To answer this question effectively, we will need to be creative and analytical.  Examination of the work site, knowledge of simultaneous operations, and competency in our job will be required.  Anticipating potential changes will insure that we are not surprised when things do change.

What will we do about it?  Again, creativity and analytical thinking are critical here.  Many times we hear the phrase, “prior planning prevents poor performance”.  Effectively answering this question insures that performance will be as designed.

How will we know we are finished?  Review of completed job hazard analysis documents has shown that it may be difficult to determine at what point the job is complete.  If the permit for the job being performed provides a scope of work like, “replace mechanical seal in hot oil pump”, once the seal is replaced, there are numerous tasks that still need to be performed before the job is complete.  Numerous times the JHA does not go beyond analyzing the tasks associated with the scope of work and do not consider additional tasks; like testing, clean up and turnover to operations.

As previously mentioned, most supervisors believe that the discussion associated with this type of analysis is more important than the completion of the form used to show that the JHA has been performed.

What about the form though?

• What happens at the conclusion of the job?
• Does anyone review the form to determine if all the hazards were found and mitigated?
• Does anyone follow up with the work group to see if anything happened that made them change the work?
• How should this review be performed and what are the benefits that will be gained by this?
• How can we learn from our experience?

Developing competent personnel is an ongoing process for most organizations.  A great deal of literature exists on the most effective methods of developing competency in adults. Training sessions are delivered using the concept of Kolb’s theory of the experiential learning cycle.  According to Kolb [2], this type of learning can be defined as “the process whereby knowledge is created through the transformation of experience.” [i] In other words, adults learn best when they are actively experiencing something and not just listening to lectures or instructor centered learning.

Experienced trainers who deliver adult learning sessions use a process of debriefing to allow reflection, reinforce learning and help the learner apply the knowledge to their life.  It is generally acknowledged in the training industry that most real learning takes place in the debrief.  This is the opportunity for learners to reflect and develop knowledge from the activity, in our case the job performed.

Very simply, debriefing a learning activity should focus on three questions.  What?  So What?  Now What?

What? is the question that guides the learning toward reflection and what just happened.  This question provides a starting point to discover what everyone involved experienced.

So What? is the question that leads to drawing conclusions and exploring alternate methods.

Now What? leads to future planning and continuous improvement initiatives that will be used to strengthen the organization’s culture and work processes.

If we return to question six of the job hazard analysis process, “How will we know we are done?”, the final answer for this question would be, “When we have completed the debrief of the job performed.”  There are five questions that should be used for debriefing a job.  These five questions, how they relate to the standard debriefing questions and the expected lessons to learn from them include:

What did we do?  This is the opportunity for reflection and to insure that the job has been completed appropriately.  Each member of the team should come to agreement that what is being described is what was actually done.   This is the What of debriefing.

Did anything change while doing the job?   Reflection on this question will lead the team to determine if the job was actually performed as it was initially described and analyzed.  This is the question that will also lead to identify incidents for investigation.  If anything unusual occurred during the task, reporting should be more efficient because the incident will be fresh in everyone’s mind.  Capturing these incidents and changes now will help modify future work orders and insure that we learn something from this experience.  This is the So What of the debriefing cycle.

Did anybody get hurt?  This question should be answered with all personnel examining themselves for strains, pulled muscles, bumps, bruises, cuts, scrapes, twisted joints, twinges in the back and a general self examination for good health.  Any small injury or potential illness should be recorded here.   This will insure that a worker does not leave the job without reporting an injury or illness, and then visit a medical provider later because something cropped up.  Having someone discover they have been injured after leaving the worksite is a problem for managers.  This allows measures to be taken early to manage the injury or illness for reporting purposes.  Here and the next question is where more exploration of the “What” is performed.

Did anybody come close to getting hurt?  This is the question that will capture near miss incidents quickly.  Near miss reporting programs fail for numerous reasons.  Lack of understanding, lack of motivation, blaming the reporter, and convenience of reporting are reasons that near misses may not be reported.  Reflection and discussion about the completed job will insure that any near miss is reported quickly.  This will lead to creation of a more comprehensive database that can be used to predict trends and identify problems areas in processes.

What would we do differently?  This is the question that will tie everything together into a plan for the future.  Recommendations and action items should be generated from this final question so that future jobs can be analyzed with more speed and efficiency.  Potential training and competency development issues may be discovered.  Procedures for modification may be identified.  Latent conditions that are not readily apparent may be identified and mitigated before they become active failures.

The Now What of the debriefing cycle is:

• Conducting an effective job task analysis and following with an effective debriefing of the job will yield several benefits.
• Competency gaps of personnel associated with the work will be identified.
• Training topics and on the job mentoring for personnel with these identified gaps, can be more quickly delivered.
• Procedural modifications that are necessary to insure that work is performed safely and efficiently will be quickly identified and addressed.
• Potential process safety incidents will be quickly identified and investigated.
• Near miss incidents will be reported quickly and the organization’s near miss/incident database will be enhanced.

The process described in this paper can be expanded to any job and any work group.  Consider an engineering team who is working on the design of a new process to be considered for construction.  Conducting an effective job task analysis in the beginning stages of the project will insure that roles, goals and expectations are addressed and known.  Conducting an effective debrief at the conclusion, or even at selected stages of a project, will enhance the project team’s effectiveness and insure that all team members are always striving to meet the goal of the project.  The attached checklist for engineering projects, at the end of this paper, may be helpful for focusing a team’s efforts.

Opportunities exist in all phases of operations and in all activities performed to keep processes safe.  It is important that all personnel be aware that learning from experience happens every day and these lessons learned need to be captured and stored for future use.

To develop process safety competency attend our PS-4, Process Safety Engineering;  HS-45, Risk Based Process Safety Management; and PS-2, Fundamental of Process Safety courses.  To develop competency in other skills, attend one of our other courses.

By Clyde Young

PetroSkills Instructor/Consultant

Reference:

1.    Young, Clyde. ,” Debrief:  The experiential learning cycle, process safety competency, safe work practices, identifying and reporting of near miss/incident data”, Mary K. O’Connor Process Safety Symposium, Texas A&M University, October 29.

2.    Kolb, David A. Experiential Learning: Experience as the Source of Learning and Development. Prentice-Hall, Inc., Englewood Cliffs, N.J. 1984.

Job Hazard Analysis Checklist

1. PROCEDURES

• ·What are the procedures for the task?
• ·What is unclear about the procedures?
• ·What order will we use these procedures?
• ·What permits are needed for hazard controls?

2. EQUIPMENT AND TOOLS

• ·What are the right tools for the job?
• ·What is the correct way to use them?
• ·What is the condition of the tool?

3. POSITIONS OF PEOPLE

• ·What could we be struck by?
• ·What could we strike ourselves against?
• ·What can we get caught in/on/between?
• ·What are potential trip/fall hazards?
• ·What are potential hand/finger pinch points?
• ·What extreme temperatures will we be in/around?
• ·What are the risks of inhaling, absorbing, swallowing hazardous substances?
• ·What are the noise levels?
• ·What electrical current/energized system could we come in contact with?
• ·What would be a cause for overexerting ourselves?

4. PERSONAL PROTECTIVE EQUIPMENT

• ·What is the proper PPE?

Hard hat, glasses/goggles, ear plugs, gloves, steel toe boots, respiratory system, fire retardant clothing

5. CHANGING THE COURSE OF WORK

• ·What would cause us to have to stop or rearrange the job?
• ·What would cause us to change our tools or equipment?
• ·What would cause us to have to change our position?
• ·What would cause us to have to change our PPE?

YOU HAVE THE RIGHT AND

THE OBLIGATION TO

STOP UNSAFE ACTS

ENGINEERING JOB ANALYSIS

1. PROCEDURES

• ·What are the procedures for the task?
• ·What is unclear about the procedures?
• ·In what order will we use these procedures?
• ·What is the proper timeline for these procedures?
• ·What permits or permissions are needed for job controls?

2. EQUIPMENT, TOOLS, DOCUMENTS

• ·What are the right tools for the job? (software, simulators, matrixes, checklists, worksheets…)
• ·What is the correct way to use them?
• ·What forms will be needed for the job?
• ·What documents will we need to produce?
• ·What else do we need to know?

3. INTERACTION WITH PEOPLE

• ·What other departments need to know about this task?
• ·Who are the personnel that need to know?
• ·What other departments will supply information for this task?
• ·Who are the personnel who will supply that information?
• ·What could prevent other personnel or departments from supplying what we need?
• ·What could prevent us from supplying what other departments need?

4.  CHANGING THE COURSE OF WORK

• ·What would cause us to have to stop or rearrange the job?
• ·What would cause us to change our tools or equipment?
• ·What would cause us to have to change our interaction with people?

YOU HAVE THE RIGHT AND THE OBLIGATION TO

STOP UNSAFE or UNPRODUCTIVE ACTS

TEG Vaporization Losses

In this Tip of The Month (TOTM), the effect of striping gas rate and triethylene glycol (TEG) circulation ratio on the TEG vaporization loss from the regenerator top and contactor top is investigated. Specifically, this study focuses on the variation of TEG vaporization losses with reboiler pressure, TEG circulation ratio and stripping gas rate. By performing a rigorous computer simulation of TEG regeneration at reboiler pressures of 110.3 kPaa (16 psia) and 524.1 kPaa (76 psia), several charts for quick estimation of TEG vaporization losses from regenerator top and contactor, which are needed for facilities type calculations are developed. In addition, the effect of contactor temperature on the TEG vaporization losses for a case study is shown.

Computer Simulation Results:

In order to study the impact of the contactor temperature, stripping gas rate and TEG circulation rate on the TEG vaporization losses, the TEG dehydration process was simulated using ProMax [1] software with its Soave-Redlich-Kwong (SRK) [2] equation of state (EOS). The process flow diagram used for these simulations is the same as in the November 2013 TOTM [3] and is shown here in  Figure 1.

The water-saturated gas with a water content of 915 mg/std m³ (57 lbm/MMSCF) enters the bottom of the contactor column at 37.8°C (100°F) and 6897 kPaa (1000 psia) at a rate of 2.835×10⁶ std m³/d (100 MMSCFD). The contactor column has three theoretical trays. The lean TEG solution enters at the top of the contactor column and flows down in the column. As shown in Figure 1, the water content of the dried gas is 10 mg/std m³ (0.63 lbm/MMSCF). The rich TEG solution contains 96.1 mass percent TEG entering the still column at 100°C (212°F) and 515 kPaa (74.7 psia). The reboiler temperature was set at 204.4°C (400°F) and boil-up ratio of 0.1 (molar bases). Two theoretical trays in the regenerator (still) column (N_R = 2) and two theoretical trays (N_S = 2) in the striping gas section were utilized. The striping gas enters the bottom of the stripping gas section at 204°C (399°F) and 524 kPaa (76 psia). Methane was used for the stripping gas at a rate of 56.3 std m³/h (1893 scf/hr). The regenerated lean solution contains 99.86 mass percent TEG and the ratio of stripping gas to lean TEG liquid volume rates is 20 std m³ of gas/std m³ of lean TEG solution (2.67 scf/sgal) or a mass ratio of 28.3. The regenerator (still) top temperature is 91.4°C  (196.5°F). If the same stripping gas was sparged directly into the reboiler (N_S = 0, no stripping gas section), with everything else remaining the same, the  regenerated solution contains 99.2 mass percent TEG and  the regenerator column top temperature remains practically the same and is 91.1°C  (196°F). For the above case the number of theoretical trays in the still column is increased from 2 to 3 (N_R = 3); the lean TEG concentration increased slightly from 99.6 to 99.8 mass percent but the regenerator column top temperature remained the same.

Using a similar set up as is shown in Figure 1, several simulations were performed for a range of stripping gas rates, for N_R=2, N_S=0 and for two reboiler pressures of 110.3 and 524 kPaa (16 and 76 psia) and temperature of 204.4°C (400°F). The results of these simulation runs are presented in Figures 2 to 5.

Figure 1. Sample results using ProMax [1] for TEG dehydration with reboiler P=110.3 kPaa (16 psia) with N_R=2 and N_S=2

The regenerator top temperatures are exactly the same as those in Figures  2 and 3 presented in the November 2013 TOTM [3].

Figure 2 presents the variation of the TEG vaporization losses from still/regeneartor column top with circulation ratio (mass basis) and stripping gas rate at top pressure of 101.3 kPaa (14.7 psia) and reboiler pressure of 110.3 kPaa (16 psia) operating at 204.4°C (400°F).

As was discussed in the November [3] and September [4] 2013 TOTMs, regeneration of TEG at higher reboiler pressure has several advantages such as preventing the emission of harmful contaminants like benzene, toluene, ethylbenzene, xylenes (BTEX), and hydrogen sulfide to the environment. Therefore, similar diagrams as shown in Figure 2 were generated for top pressure of 515.2 kPaa (74.7 psia) and reboiler pressure of 524.1 kPaa (76 psia) at 204.4°C (400°F). Figure 3 presents the variation of TEG vaporization loss from regenerator top for such a high reboiler pressure.

Fig 2. Variation of TEG vaporization loss from regenerator top with circulation mass ratio and stripping gas rate at top P=101.3 kPaa (14.7 psia) and reboiler P=110.3 kPaa (16 psia) at 204.4°C (400°F) and N_S=0

Fig 3. Variation of TEG loss from regenerator top with circulation msss ratio and stripping gas rate at top P=515.2 kPaa (74.7 psia) and reboiler P=524.1 kPaa (76 psia) at 204.4°C (400°F) and N_S=0

Figures 2 and 3 can be used for a quick estimate of the TEG vaporization loss from regenerator top for a given stripping gas rate and TEG circulation ratio either at low or high reboiler pressure. The two reboiler pressures selected in this study are typical operating pressures. For generation of data for Figures 2 and 3, the stripping gas was sparged directly into the reboiler; therefore,  the number of theoretical trays for stripping gas section is zero (N_S=0). The corresponding figures in terms of TEG circulation volume ratio are presented in the Appendix (Figures 2A and 3A). Figures 2 and 3 indicate that as the stripping gas ratio increases the TEG vaporization losses decreases. These two figures also indicate that as the TEG mass circulation ratio increases, the TEG vaporization losses increases initially followed by a decreasing trend.

Generally, either 0, 1, or 2 ideal trays in the stripping gas section is used. In order to investigate the effect of the number of theoretical trays in the stripping gas section (N_S) on the TEG vaporization loss, simulations were performed for the cases of N_S=0 and N_S=2 for a constant stripping gas rate. Figure 4 presents the results of these simulations for low reboiler presssures of  110.3 kPaa (16 psia). The reboiler temperature for all cases  was set at 204.4°C (400°F).

Figure 4 clearly indicates that the TEG vaporization loss from the regenerator top is independent of the number of ideal trays in the gas stripping section at low TEG mass circulation rates, however, it increases slightly with the increase in the number of ideal trays in the stripping gas section at higher TEG mass circulation ratio.

The effect of feed gas temperature to contactor and mass circulation ratio on TEG vaporization loss from the regenerator top is demonstrated in Figure 5. The TEG vaporization losses for three feed gas temperatures to contactor for stripping gas rate of 10 Std m³/m³ TEG (1.34 SCF/gal TEG) are plotted as a function of TEG mass circulation ratio. This figure indicates that as the contactor temperature increases, the TEG vaporization loss from the regenerator top increases. This can be explained because as the feed gas temperature increases, the feed gas water content (mass per unit volume at standard conditions) increases drastically which results in more vaporization of water from regenerator top. Consequently, more TEG (along with water) per unit volume of the gas at standard conditions is vaporized.

As expected Figure 6 indicates that the TEG vaporization loss from the contactor top is practically independent of the stripping gas rate. In addition this figure shows that as the TEG mass circulation ratio increases beyond 15 mass of TEG/mass of water removed, the TEG losses remain constant. As shown in Figure 7, Similar diagrams for higher reboiler pressure revealed almost the same observations.

Figure 8 also shows that as the TEG mass circulation ratio increases beyond 15 mass of TEG/mass of water removed, the TEG losses remain constant. However, as the feed gas temperature to the contactor increases, the TEG losses from the contactor top increase drastically.

Fig 4. Effect of number of ideal trays (N_S) in the gas stripping section on the TEG vaporization loss from regenerator top at P=101.3 kPaa (14.7 psia), reboiler P=110.3 kPaa (16 psia) at 204.4°C (400°F) and stripping gas rate of 10 Std m³/m³ TEG (1.34 SCF/gal TEG)

Fig 5. Effect of contactor temperature and mass circulation ratio on TEG vaporization loss from regenerator top at P=101.3 kPaa (14.7 psia) and reboiler P=110.3 kPaa (16 psia) at 204.4°C (400°F), N_S=0 and stripping gas rate of 10 Std m³/m³ TEG (1.34 SCF/gal TEG)

Fig 6. Variation of TEG vaporization loss from contactor top with circulation mass ratio and stripping gas rate at top P=101.3 kPaa (14.7 psia) and reboiler P=110.3 kPaa (16 psia) at 204.4°C (400°F) and N_S=0

Fig 7. Variation of TEG vaporization loss from contactor top with circulation mass ratio and stripping gas rate at top P=515.2 kPaa (74.7 psia) and reboiler P=524.1 kPaa (76 psia) at 204.4°C (400°F) and N_S=0

Fig 8. Effect of contactor temperature and mass circulation ratio on TEG vaporization loss from contactor top at P=6897 kPaa (1000 psia), reboiler P=110 kPaa (16 psia) at 204.4°C (400°F) , N_S=0 and stripping gas rate of 10 Std m³/m³ TEG (1.34 SCF/gal TEG)

Conclusions:

In this TOTM, the effect of circulation ratio, stripping gas rate, theoretical number of trays, and the feed gas temperature to the contactor column on the TEG vaporization losses from contactor  top and regenerator top for regeneration of TEG concentration at low and high reboiler pressure operating at 204.4°C (400°F) was studied. Charts for a quick estimation of the TEG vaporization losses from still/regenerator column top and contactor top at a specified stripping gas rate and circulation ratio to achieve a desired level of lean TEG concentration were prepared and presented in Figures 2, 3, 6, and 7. These charts are based on the rigorous calculations performed by computer simulations and can be used for facilities type calculations for evaluation and trouble shooting of an operating TEG dehydration unit. In addition, the following observations were made for the cases studied in this TOTM:

1. The TEG vaporization loss from the contactor top is almost 10 times higher than still/regenerator column top (see Figures 2, 3, 6 and 7).
2. As the feed gas temperature to the contactor column increases, the TEG vaporization loss from top of both columns increases (Figures 5 and 8).
3. The TEG vaporization loss from top of still/regenerator column is practically independent of the number of theoretical trays in the stripping gas section (N_S), see Figure 4.
4. Pressurized reboiler results in higher TEG vaporization losses from regenerator due to higher stripping gas requirements.
5. Even though not studied in this TOTM, mechanical losses such as entrainment from contactor top and regenerator top as well as leaks from pump seals are much higher than the vaporization losses presented here.

To learn more, we suggest attending our G40 (Process/Facility Fundamentals), G4 (Gas Conditioning and Processing), G5 (Gas Conditioning and Processing-Special), and PF81 (CO₂ Surface Facilities), PF4 (Oil Production and Processing Facilities), courses.

John M. Campbell Consulting (JMCC) offers consulting expertise on this subject and many others. For more information about the services JMCC provides, visit our website at www.jmcampbellconsulting.com, or email us at consulting@jmcampbell.com.

By: Dr. Mahmood Moshfeghian

References:

1. ProMax 3.2, Bryan Research and Engineering, Inc., Bryan, Texas, 2013.
2. Soave, G., Chem. Eng. Sci. Vol. 27, No. 6, p. 1197, 1972.
3. Moshfeghian, M., http://www.jmcampbell.com/tip-of-the-month/2013/11/estimating-still-column-top-temperature-in-teg-dehydration-unit/, Tip of the Month, November 2013.
4. Moshfeghian, M., http://www.jmcampbell.com/tip-of-the-month/2013/09/high-pressure-regeneration-of-teg-with-stripping-gas/, Tip of the Month, September 2013.

Appendix

Fig 2A. Variation of TEG vaporization loss from regenerator top with circulation volume ratio and stripping gas rate at top P=101.3 kPaa (14.7 psia) and reboiler P=110.3 kPaa (16 psia) at 204.4°C (400°F) and N_S=0

Fig 3A. Variation of TEG loss from regeneartor top with circulation volume ratio and stripping gas rate at top P=515.2 kPaa (74.7 psia) and reboiler P=524.1 kPaa (76 psia) at 204.4°C (400°F) and N_S=0

In this Tip of The Month (TOTM), the effect of striping gas rate and TEG circulation ratio on the still column top temperature for regeneration of rich triethylene glycol (TEG) is investigated. Specifically, this study focuses on the variation of still column top temperature with reboiler pressure, TEG circulation ratio and stripping gas rate. By performing a rigorous computer simulation of TEG regeneration at reboiler pressures of 110.3 kPaa (16 psia) and 524.1 kPaa (76 psia), two charts for quick determination of still column top temperature needed for facilities type calculations are developed. In addition, the effect of theoretical number of trays in the stripping gas section is studied.

Computer Simulation Results:

In order to study the impact of stripping gas rate and TEG circulation rate on the still column top temperature, the TEG dehydration process was simulated using ProMax [1] software with its Soave-Redlich-Kwong (SRK) [2] equation of state (EOS). The process flow diagram used for these simulations is shown in  Figure 1.

The water-saturated gas with a water content of 915 mg/std m³ (57 lbm/MMSCF) enters the bottom of the contactor column at 37.8°C (100°F) and 6895 kPaa (1000 psia) at a rate of 2.835×10⁶ std m³/d (100 MMSCFD). The contactor column has three theoretical trays. The lean TEG solution enters at the top of the contactor column and flows down in the column. As shown in Figure 1, the water content of the dried gas is 10 mg/std m³ (0.63 lbm/MMSCF). The rich TEG solution contains 96.1 mass percent TEG entering the still column at 100°C (212°F) and 515 kPaa (74.7 psia). The reboiler temperature was set at 204.4°C (400°F) and boil-up ratio of 0.1 (molar bases). Two theoretical trays in the regenerator (still) column (N_R = 2) and two theoretical trays (N_S = 2) in the striping gas section were specified. The striping gas enters the bottom of the stripping gas section at 204°C (399°F) and 524 kPaa (76 psia). Methane was used for the stripping gas at a rate of 56.3 std m³/h (1893 scf/hr). The regenerated lean solution contains 99.6 mass percent TEG and the ratio of stripping gas to lean TEG liquid volume rates is 20 std m³ of gas/std m³ of lean TEG solution (2.67 scf/sgal) or a mass ratio of 28.3. The regenerator (still) top temperature is 91.4°C  (196.5°F). If the same stripping gas was sparged directly into the reboiler (N_S = 0, no stripping gas section), with everything else remaining the same, the  regenerated solution contains 99.2 mass percent TEG and  the regenerator column top temperature remains practically the same and is 91.1°C  (196°F). For the above case the number of theoretical trays in the still column is increased from 2 to 3 (N_R = 3); the lean TEG concentration increased slightly from 99.6 to 99.8 mass percent but the regenerator column top temperature remained the same.

Using a similar set up as is shown in Figure 1, several simulations were performed for a range of stripping gas rates, for N_R=2, N_S=0 and for two reboiler pressures of 110.3 and 524 kPaa (16 and 76 psia) and temperature of 204.4°C (400°F). The results of these simulation runs are presented in Figures 2 to 5.

Figure 1. Sample results using ProMax [1] for TEG dehydration with reboiler P=110.3 kPaa (16 psia) with N_R=2 and N_S=2

Figures 2 presents the variation of still column top temperature with circulation ratio (mass basis) and stripping gas rate at top pressure of 101.3 kPaa (14.7 psia) and reboiler pressure of 110.3 kPaa (16 psia) operating at 204.4°C (400°F).

As was discussed in the August 2013 TOTM, regeneration of TEG at higher reboiler pressure has several advantages such as preventing the emission of harmful contaminants like benzene, toluene, ethylbenzene, xylenes (BTEX), and hydrogen sulfide to the environment [3]. Therefore, similar diagrams as shown in Figure 2 were generated for top pressure of 515.2 kPaa (74.7 psia) and reboiler pressure of 524.1 kPaa (76 psia) at 204.4°C (400°F). Figure 3 presents the variation of still column top temperature for such a high reboiler pressure.

Fig 2. Variation of still column top temperature with circulation mass ratio and stripping gas rate at top P=101.3 kPaa (14.7 psia) and reboiler P=110.3 kPaa (16 psia) at 204.4°C (400°F)

Fig 3. Variation of still column top temperature with circulation msss ratio and stripping gas rate at top P=515.2 kPaa (74.7 psia) and reboiler P=524.1 kPaa (76 psia) at 204.4°C (400°F)

Figures 2 and 3 can be used for a quick determination of the still column top temperature for a given stripping gas rate and TEG circulation ratio either at low or high reboiler pressure. The two reboiler pressures selected in this study are typical operating pressures. For generation of data for Figures 2 and 3, the stripping gas was sparged directly into the reboiler; therefore,  the number of theoretical trays for stripping gas section is zero (N_S=0). The corresponding figures in terms of TEG circulation volume ratio are presented in the Appendix (Figures 2A and 3A).

Generally, either 0, 1, or 2 theoretical trays in the stripping gas section is used. In order to investigate the effect of the number of theoretical trays in the stripping gas section (N_S) on the still column top temperature, simulations were performed for the cases of N_S=0 and N_S=2 for two constant stripping gas rates.

Figures 4 and 5 present the results of these simulations for low and high reboiler presssures of  110.3 kPaa (16 psia) and 524.1 kPaa (76 psia), respectively. The reboiler temperature for all cases  was set at 204.4°C (400°F).

Figures 4 and 5 clearly indicate that the still column top temperature is independent of the number of theoretical trays in the stripping gas section. Therefore, Figures 2 and 3 can be used for any number of theoretical trays in the stripping gas section.

Fig 4. Effect of the number of theoretical trays (N_S) on the still column top temperature at various circulation ratio and stripping gas rate at top P=101.3 kPaa (14.7 psia) and reboiler P=110.3 kPaa (16 psia) at 204.4°C (400°F)

Fig 5. Effect of the number of theoretical trays (N_S) on the still column top temperature at various circulation ratio and stripping gas rates at top P=515.2 kPaa (74.7 psia) and reboiler P=524.1 kPaa (76 psia) at 204.4°C (400°F)

Similar study also showed that the feed gas temperature to the contactor column has no effect on the still column top temperature. The results of this study are shown in Figures 6 and 7 of the Appendix.

Conclusions:

In this TOTM, the effect of circulation ratio, stripping gas rate, theoretical number of trays, and the feed gas temperature to the contactor column on the still column top temperature for regeneration of TEG concentration at low and high reboiler pressure operating at 204.4°C (400°F) was studied. Two charts for a quick determination of the still column top temperature at a specified stripping gas rate and circulation ratio to achieve a desired level of lean TEG concentration were prepared and presented in Figures 2 through 3 (see the corresponding figures in the Appendix). These charts are based on the rigorous calculations performed by computer simulations and can be used for facilities type calculations for evaluation and trouble shooting of an operating TEG dehydration unit. In addition, the following observations were made:

1. The still column top temperature is independent of the number of theoretical trays in the stripping gas section (N_S) and feed gas temperature to the contactor column.
2. As the stripping gas rate increased, the still column top temperature decreased.
3. As the TEG circulation ratio increased, the still column top temperature decreased.
4. Pressurized reboiler results in much higher still column top temperature than the atmospheric reboiler.

To learn more, we suggest attending our G40 (Process/Facility Fundamentals), G4 (Gas Conditioning and Processing), G5 (Gas Conditioning and Processing-Special), and PF81 (CO₂ Surface Facilities), PF4 (Oil Production and Processing Facilities), courses.

John M. Campbell Consulting (JMCC) offers consulting expertise on this subject and many others. For more information about the services JMCC provides, visit our website at www.jmcampbellconsulting.com, or email us at consulting@jmcampbell.com.

By: Dr. Mahmood Moshfeghian 

References:

1. ProMax 3.2, Bryan Research and Engineering, Inc., Bryan, Texas, 2013.
2. Soave, G., Chem. Eng. Sci. Vol. 27, No. 6, p. 1197, 1972.

Moshfeghian, M., http://www.jmcampbell.com/tip-of-the-month/2013/08/teg-dehydration-how-does-the-stripping-gas-work-in-lean-teg-regeneration/, Tip of the Month, August 2013.

Appendix

Fig 2A. Variation of still column top temperature with circulation volume ratio and stripping gas rate at top P=101.3 kPaa (14.7 psia) and reboiler P=110.3 kPaa (16 psia) at 204.4°C (400°F)

Fig 3A. Variation of still column top temperature with circulation volume ratio and stripping gas rate at top P=515.2 kPaa (74.7 psia) and reboiler P=524.1 kPaa (76 psia) at 204.4°C (400°F)

Fig 6. Variation of still column top temperature with circulation mass ratio and feed gas temperature to the contactor column at a specified stripping gas rate at top P=101.3 kPaa (14.7 psia) and reboiler P=110.3 kPaa (16 psia) at 204.4°C (400°F)

Fig 7. Variation of still column top temperature with circulation mass ratio and feed gas temperature to the contactor column at a specified stripping gas rate at top P=515.2 kPaa (74.7 psia) and reboiler P=524.1 kPaa (76 psia) at 204.4°C (400°F)

Stainless steel is a family of corrosion resistant steels containing chromium in which the chromium forms a passive film of chromium oxide (Cr₂O₃) when exposed to oxygen [1]. This phenomenon is called passivation and is seen in other metals, such as aluminium and titanium.

The film layer is impervious to water and air and quickly reforms when the surface is scratched. This protects the metal beneath – preventing further surface corrosion.  Since the layer only forms in the presence of oxygen, corrosion-resistance can be adversely affected if the component is used in a non-oxygenated environment e.g. underwater bolts on a platform support structure.

Such passivation only occurs if the proportion of chromium is high enough and is normally achieved with addition of at least 13% (by weight) chromium. Progressively higher levels of corrosion resistance and strength is achieved  by the addition of other alloying elements – each offering specific attributes in respect of strength and corrosion resistance.

Classification issues

The need to classify stainless steel has led to a fundamental problem of which method to use. Probably the best known system derives from of the Society of Automobile Engineers (SAE) e.g. 316 Cr/Ni/Mo 17/12/2. This is interpreted as stainless steel containing the proportions of 17% chromium, 12% nickel, and 2% molybdenum.

However, the waters are somewhat muddied by a variety of international and country-based systems that include EN (European Norm); and UNS (Unified Numbering System). For example, SAE 304 Cr/Ni 18/10 stainless steel is EN 1.4301 which is UNS S30400.

Stainless steels may also be graded into five basic families or phases determined by their crystalline structure: the stable phases austenitic or ferritic; a duplex mix of the two; the martensitic phase created when some steels are quenched from a high temperature; and precipitation-hardenable.

Ferritic stainless steel

In ferritic stainless steel, the iron and chromium atoms are arranged in what is termed a body-centred cubic structure in which the atoms are arranged on the corners of the cube and one in the centre (Figure 1). As well as being ferro-magnetic, ferritic stainless steel exhibits very high stress corrosion cracking resistance.

Ferritic stainless steels are plain chromium (10.5 to 18%) grades characterized by moderate corrosion resistance and poor fabrication properties. These characteristics may be improved with the addition of molybdenum; some, aluminium or titanium.

Austenitic stainless steel

With the addition of nickel, the properties change dramatically. As shown (Figure 3) the atoms are re-arranged so that they occur on the corners of the cube and also in the centre of each of the faces. In this manner it becomes what is termed austenitic stainless steel.

It can thus be seen from Table 1 that unless you are specifically looking for a ferro-magnetic material, austenitic stainless steel would be the most obvious choice. Indeed this is borne out by the fact that austenitic stainless steels account for about 70% or more of all stainless steel used worldwide – with ferritic stainless steels making up about 25%. The other families each represent less than 1% of the total market.

Austenitic stainless steels are designated by numbers in the 200 and 300 series.

Series 300

The relationship between the 300 austenitic grades is shown in Figures 4.

The basic grade 304 contains about 18% chromium and 8% nickel (often referred to as 18/8) and range through to the high alloy or ‘super austenitics’ such as 904L and 6% molybdenum grades.

Additional elements can be added such as molybdenum, titanium or copper, to modify or improve their properties, making them suitable for many critical applications involving high temperature as well as corrosion resistance. This group of steels is also suitable for cryogenic applications because the effect of the nickel content in making the steel austenitic avoids the problems of brittleness at low temperatures, which is a characteristic of other types of steel.

Generally, the 300 grade alloys are subject to crevice and pitting corrosion.

Low-carbon versions, (indicated by the letter suffix L) include 304L, 316L and 317L, in which the carbon content of the alloy is below 0.03%. This reduces the effect of ‘sensitisation’ in which chromium carbides precipitate at the grain boundaries due to the high temperatures involved in welding.  The relatively high nickel content also inhibits the brittleness exhibited by ferritic materials at low temperatures and thus makes austenitic steels suitable for cryogenic applications.

200 series

We have seen earlier how the addition of nickel is used in the creation of the classic chrome-nickel 300 series austenitic stainless steel.

The reduced nickel content of the 200 series chrome-manganese grades makes them significantly cheaper. However, depending on their chemistry, they also offer good formability (ductility) and/or strength.  Indeed, certain grades (201, 202 and 205 series) even offer about 30% higher yield strength than the classic 304-series chrome-nickel grade – allowing designers to cut weight (Table C.2).

Reducing nickel, on the other hand, reduces the maximum chromium content possible in the alloy. Less chromium means less corrosion resistance and a consequent narrowing of the range of applications for which the material is suitable.

A word of warning comes from the International Stainless Steel Forum (ISSF). Continuous pressure to cut costs, especially from the Asian market, has resulted in the development of austenitic grades ever lower in nickel and chromium, often not covered by international codes or specifications. In fact, numerous chrome-manganese grades are company-specific and identified simply by a title given to them by the producer.

Duplex stainless steels

Duplex stainless steels [6] are a mixture of austenite and ferrite microstructures that combine some of the features of each class:

• resistance to stress corrosion cracking  – but inferior to ferritic steel;
• superior toughness to ferritic steel – but inferior to austenitic steel;
• roughly twice the strength of austenitic steel;
• superior resistance to pitting, crevice corrosion and stress corrosion cracking;
• high resistance to chloride ions attack; and
• high weldability.

These features are achieved by adding less nickel than would be necessary for making a fully austenitic stainless steel. Thus, Grade 2304 comprises 23% chromium and 4% nickel whilst Grade 2205 comprises 22% chromium and 5% nickel – with both grades containing further minor alloying additions.

On the negative side, austenitic-ferritic duplex stainless steels are only usable between temperature limits of about -50°C and 300°C – outside which they suffer reduced toughness.

Martensitic stainless steel

Named after the German metallurgist, Adolf Martens, the martensitic Grade 400 series (Figure 5) are low carbon (0.1–1%), low nickel (less than 2%) steels containing chromium (12 to 14%) and molybdenum (0.2–1%).

Stainless steels hardened by transformation to martensite are tempered to give the desired engineering properties. At high temperatures they have an austenitic structure that is transformed into martensitic structure upon cooling to room temperature. Unfortunately, this tempering can influence corrosion susceptibility. For example, corrosion susceptibility of type 420 stainless steel is at its maximum when the alloy is tempered at temperatures in the range of 450° to 600°C. So, aalthough not as corrosion-resistant as the 200 and 300 classes, martensitic stainless steels are magnetic, extremely strong (if not a little brittle), highly machinable, and can be hardened by heat treatment.

Martensitic stainless steels are subject to both uniform and non-uniform attack in seawater. And the incubation time for non-uniform attack in even weak chlorides is often only a few hours or days.

Precipitation-hardening martensitic stainless steels

These chromium-  and nickel-containing steels can be precipitation hardened to develop very high tensile strengths. Precipitation-hardening stainless steels are usually designated by a trade name rather than by their AISI 600 series designations.

The most common grade in this group is ‘17-4 PH’, also known as Grade 630, with a composition of 17% chromium, 4% nickel, 4% copper and 0.3% niobium. The main advantage of these steels is that they can be supplied in the ‘solution treated’ condition – in which state the steel is just machinable. Following machining, forming, etc. the steel can be hardened by a single, fairly low temperature ‘ageing’ heat treatment that causes no distortion of the component.

Precipitation hardening generally results in a slight increase in corrosion susceptibility and an increased susceptibility to hydrogen embrittlement.

By: Mick Crabtree

References

1. T. Sourmail and H. K. D. H. Bhadeshia, ‘Stainless Steels’, University of Cambridge.

2. Nabil Al-Khirdaji, ‘Stainless Steel Family’, Kappa Associates International.

3. ‘Stainless Steel and Corrosion’, ArcelorMittal, Stainless Europe.

4. A.U. Malik, M. Kutty, Nadeem Ahmad Siddiqi, Ismaeel N. Andijani, and Shahreer Ahmad, ‘Corrosion Studies on SS 316 L in low pH high Chloride product water medium’, 1990.

5. ‘The Stainless Steel Family’, International Stainless Steel Forum.

6. API Technical Report 938-C: ‘Use of Duplex Stainless Steels in the Oil Refining Industry’.

In this Tip of The Month (TOTM), regeneration of rich triethylene glycol (TEG) with striping gas at high pressure is investigated. Specifically, this study focusses on the determination of the required stripping gas rate as a function of the lean TEG mass percent, reboiler temperature, and the number of theoretical trays in the stripping section (N_S) for a regenerator (still) column with two theoretical trays (N_R). By performing rigorous computer simulations of TEG regeneration at high pressure, a series of charts for quick determination of stripping gas rates needed for facilities type calculations are developed. The results of this study are also compared with the case of TEG regeneration at atmospheric pressure, which was published in the June 2013 TOTM.

In gas dehydration service, TEG absorbs limited quantities of benzene, toluene, ethylbenzene, and xylene ( BTEX) along with other volatile organic compounds (VOCs)  from the gas. BTEX and other VOC emissions are an environmental challenge for the natural gas industry since some of these compounds are considered to be carcinogenic.

Absorption is fa­vored at lower temperatures, higher pressure, increased lean TEG concentration and circulation rate. Based on the data from the June 2011 TOTM [1], predicted absorption levels for BTEX components vary from 5 to 15% for benzene, 5 to 25% for toluene and ethylbenzene, and 5 to 35% for xylene [2]. Some of these absorbed BTEX components are released with the flashed gas off of the TEG separator operating at around 483 kPag (70 psig). The flashed gas may be used as fuel or sent to the flare. The rich TEG solution is normally regenerated at low pressure and high temperature. The remaining BETX components are released with the vaporized water and stripping gas at the top of the still column. Hicks et al. [3] discuss different options for VOC emission control. One of their options was to operate the reboiler at pressures that allow the overhead vapors that contain VOCs including BTEX and stripping gas to flow directly to either compressor suction, a fuel system, or to a flare. Their study showed higher pressure reboilers effectively control BTEX emissions and economically recover all the hydrocarbon vapors with minimal incremental capital cost and no required emissions monitoring. Hicks et al. [3] highlighted the following advantages for the pressurized glycol regeneration system:

• All vapors (VOCs, H₂S, CO₂) released from the glycol are at a sufficient pressure, which allows the use of simple handling methods without discharge to the atmosphere. Typical handling methods are:

1. Mixing the released vapors with fuel gas or other gas users.
2. Using compressors in other service instead of dedicated compression to recompress the released vapors.
3. Using dedicated compression which allows for better reboiler pressure operation.

• Because the previously mentioned methods can handle vapors from the still overheads, one can eliminate the equipment specifically designed to recover these vapors. This equipment typically includes condensers, separators, pumps, and a means for burning or incinerating the non-condensable vapors. The entire still column overhead can be routed to the compression suction scrubbers with no intervening components like a condenser or separator.
• The larger-than-normal amount of stripping gas used in the stripper results in less VOCs in the condensed water vapor from the overheads. This is because the vapor-liquid equilibrium is shifted such that condensation of VOCs is minimized. The disposal water will therefore contain significantly less undesirable components. Similarly, for those systems in which the overheads are mixed with other gases, the volume ratio of overhead to mixing gas (such as gases for fuel and compressor suction) will further reduce the VOCs in the condensed water.
• No environmental testing for atmospheric discharge of VOCs is required because no VOCs will be released, as is often the case with conventional technology. Furthermore, discharge of VOCs from the regeneration system at considerably higher-than-normal pressures facilitates mixing and extreme dilution of VOCs in plant fuel or in fuel sold and transmitted to remote users.

Computer Simulation Results: 

In order to study the impact of stripping gas rate on the lean TEG mass percent, the TEG regeneration process was simulated using ProMax [4] software with its Soave-Redlich-Kwong (SRK) [5] equation of state (EOS). The process flow diagram used for these simulations is shown in  Figure 1.

Figure 1. Sample results using ProMax [3] for high pressure TEG regeneration at reboiler P=515 kPaa (74.7 psia) and N_S=2

As shown in Figure 1, the rich TEG solution contained 97.5 mass percent TEG entering the still column at 150°C (302°F) and 515 kPaa (74.7 psia). The reboiler temperature was set at 204°C (400°F) and boil-up ratio of 0.1 (molar bases).  Two theoretical trays in the still column (N_R = 2) and two theoretical trays (N_S = 2) in the gas striping section were specified. The striping gas enters the bottom of the gas stripping section at 150°C (302°F) and 521.9 kPaa (75.7 psia). Methane was used for the stripping gas at a rate of 70 std m³/h (2458 scf/hr). The regenerated lean solution contains 99.25 mass percent TEG and the ratio of stripping gas to lean TEG liquid volume rates is 80.26 std m³ of gas/std m³ of lean TEG solution (10.73 scf/sgal). If stripping gas was sparged directly into the reboiler (N_S = 0, no gas stripping section), with everything else remaining the same, the  regenerated solution contains 98.49 mass percent TEG and  the ratio of stripping gas to lean TEG liquid volume rates is 79.58 std m³ of gas/std m³ of lean TEG solution (10.64 scf/sgal). For the above cases the number of theoretical trays in the still column is increased from 2 to 3 (N_R = 3) and the lean TEG concentration remained almost the same. The concentration of rich TEG solution is also varied from 90 to 98 mass percent, but no appreciable change in the lean TEG concentration was observed for the same stripping gas rate.

Using a similar set up as is shown in Figure 1, several simulations were performed for; a range of stripping gas rates, for N_R=2, N_S=0, 1, 2 and for reboiler temperatures of 204, 193, and 182°C (400, 380, and 360°F) and a reboiler pressure of 515 kPaa (74.7 psig). The results of these simulation runs are presented in Figures 2 to 4. For purposes of comparison all of these diagrams are replotted in Figure 5.

Fig 2. Effect of lean TEG mass %, reboiler temperature and number of theoretical trays in stripping gas column (N_S=0) at reboiler P=515 kPaa (74.7 psia) on the stripping gas rate

Fig 3. Effect of lean TEG mass %, reboiler temperature and number of theoretical trays in stripping gas column (N_S=1) at reboiler P=515 kPaa (74.7 psia)  on the stripping gas rate

Fig 4. Effect of lean TEG mass %, reboiler temperature and number of theoretical trays in stripping gas column (N_S=2) at reboiler P=515 kPaa (74.7 psia)  on the stripping gas rate

Figure 6 presents the required stripping gas rate for both a pressurized reboiler, 515 kPaa (74.7 psia), and a low pressure, 109 kPaa (15.8 psia) reboiler having two theoretical trays (N_S=2) in the gas stripping section. This figure indicates that required stripping gas rates are 10 to 100 times higher for pressurized reboiler when compared to the low pressure reboiler.

Conclusions:

In this TOTM, the effect of stripping gas rate on the regenerated lean TEG concentration at high pressure for several operation conditions was studied. A series of charts to use for a quick determination of the required stripping gas rate to achieve a desired level of lean TEG concentration are presented in Figures 2 through 6. These charts are based on the rigorous calculations performed by computer simulations and can be used for facilities type calculations for evaluation and trouble shooting an operating TEG dehydration unit. In addition, the following observations were made:

1. For rich TEG concentrations between 90 and 98 mass percent, the required stripping gas rate is independent of rich TEG concentration.
2. As the number of theoretical trays in the stripping column (N_S) increases from 0 to 2, the required striping gas rate decreases.
3. Increasing the number of theoretical trays in the still column (N_R) from 2 to 3 has no appreciable effect on the stripping gas requirement.
4. Increasing the reboiler temperature from 182 to 204 °C (360 to 400 °F), decreases the required stripping gas rate.
5. Increasing the TEG reboiler pressure from 109 kPaa (15.8 psia) to 515 kPaa (74.7 psia) increases the stripping gas requirement by a factor of between 10 and 100 depending on other factors.

Fig 5. Effect of lean TEG mass %, reboiler temperature and number of ideal trays in stripping gas column on the stripping gas requirement at reboiler P=515 kPaa (74.7 psia).

To learn more, we suggest attending our G40 (Process/Facility Fundamentals), G4 (Gas Conditioning and Processing), G5 (Gas Conditioning and Processing-Special), and PF81 (CO₂ Surface Facilities), PF4 (Oil Production and Processing Facilities), courses.

John M. Campbell Consulting (JMCC) offers consulting expertise on this subject and many others. For more information about the services JMCC provides, visit our website at www.jmcampbellconsulting.com, or email us at consulting@jmcampbell.com.

By: Dr. Mahmood Moshfeghian

Fig 6. Comparison of high pressure with low pressure stripping gas requirements for TEG regeneration. HP=515 kPaa (74.7 psia), LP=109 kPaa (15.8 psia) with N_S=2

References:

1. Moshfeghian, M., http://www.jmcampbell.com/tip-of-the-month/2011/06/absorption-of-aromatics-compounds-in-teg-dehydration-process/, Tip of the Month, June 2011.
2. Moshfeghian, M. and R.A. Hubbard, “Quick Estimation of Absorption of Aromatics Compounds (BTEX) in TEG Dehydration Process,” Proceedngs of the 3rd International Gas Processing Symposium, March 5 – 7 2012 , Qatar , 2012.
3. Hicks, R., Gallaher, D. and R. Craig, “Pressurized reboiler reduces VOC emissions in glycol dehy systems”, Oil & gas j., Vol 102, Issue 17, April 2004.
4. ProMax 3.2, Bryan Research and Engineering, Inc., Bryan, Texas, 2013.
5. Soave, G., Chem. Eng. Sci. Vol. 27, No. 6, p. 1197, 1972.