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Dynamic analysis and optimization of flare network system for topside process of offshore plant
Journal Pre-proof Dynamic analysis and optimization of flare network system for topside process of offshore plant Yeon-pyeong Jo, Yongheon Cho, Sungwon Hwang
PII:
S0957-5820(19)31781-1
DOI:
https://doi.org/10.1016/j.psep.2019.12.008
Reference:
PSEP 2022
To appear in:
Process Safety and Environmental Protection
Received Date:
8 September 2019
Revised Date:
18 November 2019
Accepted Date:
9 December 2019
Please cite this article as: Jo Y-pyeong, Cho Y, Hwang S, Dynamic analysis and optimization of flare network system for topside process of offshore plant, Process Safety and Environmental Protection (2019), doi: https://doi.org/10.1016/j.psep.2019.12.008
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Dynamic analysis and optimization of flare network
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Yeon-pyeong Jo, Yongheon Cho and Sungwon Hwang*
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system for topside process of offshore plant
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Department of Chemistry and Chemical Engineering, Inha University, Incheon 22212, South
Corresponding Author
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Korea
Highlights
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* E-mail address: [email protected].
Design and optimization of flare relief system in emergency situation at offshore plant
Designing flare relief system at steady-state has an excessive margin for line size
Reduction of overall capital cost is achieved by using dynamic simulation
Enhancement of flare system safety considering uneven distribution of flare loads
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ABSTRACT
This study introduces a new approach for designing a flare network system that ensures economic feasibility and safety using dynamic simulation analysis with the gPROMS ProcessBuilder software. A case of “separator outlet blocked discharge” was selected as a
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pressure-relief scenario based on the American Petroleum Institute Standard 521, and design data from a previous offshore project were used. As the main process and flare network system were
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dynamically simulated, the effects of the pressure, temperature, and liquid-level changes in the vessel on the opening of the pressure safety valve (PSV) were analyzed. Then, dynamic
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simulation results, including those for the flare load, PSV back pressure, and Mach number, were
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compared with those obtained from a steady-state model employing the Aspen Flare System Analyzer. Lastly, the sizes of the PSVs, branch lines, and main headers were optimized to
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minimize the overall capital costs and ensure the safety of the flare network system. This methodology can be applied to all existing and newly designed flare network system to enhance
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safety and reduce capital costs.
Keywords: Emergency flare emission, Dynamic simulation, Mach number calculation, Line
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sizing criteria, Pipe line optimization
1.
Introduction
In large-scale chemical plants, various types of equipment operate, mainly including rotating
machines, separators, reactors, heat exchangers, and vessels. They normally run under the allowable operating conditions (e.g., temperatures, pressures, and flowrates) specified by either the manufactures or the operators. However, in emergency situations, such as fires, utility
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failures, and outlet blockages, such equipment tends to operate abnormally, leading to overpressure of the equipment (American Petroleum Industry, 2014). This may result in serious fires and explosions, causing severe damage to humans, property, and the environment (Shin et al., 2018). To prevent such accidents, pressure-relief systems are employed in most chemical plants. This pressure-relief systems protect the process equipment and relieve overpressure that
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arises during normal or abnormal operation. Furthermore, the emergency shutdown (ESD) system is widely used to protect the equipment
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and humans. Pengyou Zhu (Zhu et al., 2019) developed an overall framework to identify gaps
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and challenges of ESD system based on the field study in Norway. The authors mentioned that the workflow of ESD system includes three key parameters (i.e., technical, operational,
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organizational parameters). The technical parameter is related to equipment and systems, the organizational parameter to personnel with defined roles or functions and specific competence,
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and the operational parameter to actions and activities that the personnel has to perform. According to identification results using the framework, the authors found 15 challenges during
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the operation and maintenance of ESD system. Among the 15 challenges, three, seven and five of them were related to technical, organizational, and operational challenges, respectively. As a result, the paper provided effective guideline regarding the operation and maintenance of ESD
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system.
In addition, the emergency situation can be handled by applying an integrated robust and
resilient control scheme to the process. In such a case, the role of robust control is to guarantee the product specifications in emergency situation, and the role of resilient control is to stabilize the process fast during an emergency situation (Jin et al, 2010). However, further research is
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necessary, and it is important that the flare network system should be integrated with such control system. Moreover, in terms of risk assessment, the emergency scenario can be predicted based on historical data. Faisal Khan (Khan et al., 2015) presented the overall risk assessment method. In the study, the authors adopted various methods to manage the risk of the process. One of the
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methods, which used the historical data, was to combine the Fault Tree Analysis (FTA) with the Event Tree Analysis (ETA). FTA graphically illustrated the logical relationship between cause,
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consequence and failure paths. ETA evaluated the sequence of failure processes and analyzed its
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consequences. The authors described System Hazard Identification, Prediction and Prevention (SHIPP), which combined FTA and ETA. SHIPP is able to predict the probability of emergency
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by analyzing the history of emergency. As a results, the frequency of emergency situations can be predicted.
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Figure 1 shows the simplified structure of the pressure-relief system. It contains pressure-relief devices, headers and tail pipes that transfer flare loads, knock-out drums that separate liquid and
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vapor before combustion, and a flare stack that burns flare gas safely (Bader et al., 2011). Pressure-relief devices typically include a pressure safety valve (PSV), a blowdown valve (BDV), and a rupture disk, and they release fluid from the inside of the equipment to relieve
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overpressure. The PSV performs depressurization automatically under abnormal conditions that cannot be recognized by the operator. It normally opens via spring or elastic action when the pressure of the equipment reaches a designated set pressure. The BDV also performs depressurization, but it is generally used for manual operation to release fluid quickly from the inside of the vessel for process shutdown or equipment replacement. While the PSV constantly maintains the pressure of the vessel below the set pressure, the BDV reduces the pressure of the
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vessel to the level set by the operator (Bjeere et al, 2016). A rupture disc is installed in facilities or devices, such as pressure containers, pipes, and reactors, and it is ruptured when the pressure reaches the set point during operation to protect the equipment. Thus, it is only used once and should be replaced with a new one once it is ruptured. In general, it is used as a spare device of the PSV when the valve function is impaired because abnormal materials have accumulated in
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the PSV and the valve cannot respond quickly enough. The most commonly used relief device in emergency situations is the PSV (Roberts et al, 2004). The design of the PSV depends on the
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size of the orifice, the flare load, and its back pressure. Therefore, it is important to install PSVs
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of the correct type and size, as well as the correct number of PSVs, for safe depressurization.
Figure 1. Typical structure of the flare relief system
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Flare gas released from each pressure-relief device, which is installed in various equipment, flows through a tail pipe and a sub-header and gathers in the main header. Subsequently, the relieved vapor and liquid are collected in a knockout drum (K.O Drum). Then, gas is eventually sent to the flare stack for incineration before its release to the atmosphere (Lee et al., 2012). During normal operation of the flare system, when there is no released fluid from the PSV, small
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amounts of light hydrocarbon vapor flow through the header to maintain the operation of the flare network (Chopinet and Marcec-Rahelic, 2015). However, in the case of an emergency or
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abnormal operation, large amounts of released fluids flow through the header with a high
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velocity at a high pressure (Song et al., 2014). In general, the following aspects should be
Limits of the PSV back pressure
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Fluid Mach number
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Momentum of the released fluid
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considered when engineers design flare pipes (i.e., tail pipes, sub-headers, and header pipes).
Generally, the engineers consider the most governing PSV relieving scenarios when they
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design the flare system. For example, all feasible abnormal scenarios, such as total power failure, external fires, and equipment failure, are considered for designing the size of the flare pipes, size of the PSV orifices, number of PSVs, etc. Among them, the most governing case is considered
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for designing these parameters of the flare system. Furthermore, an additional margin is added to the developed design parameters to cover unanticipated emergency situations. This leads to increased capital costs of the equipment, limited space for equipment installation, and additional weight of the offshore plant (Goyal and Al-Ansari, 2009). Over the past few decades, many researchers have investigated the design of the flare system. For example, Muktikana Sahoo (Sahoo, 2013) developed a steady-state model of the flare system
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in an external fire scenario using industry data and estimated the PSV back pressure and velocity of the relieving fluid. The results indicated that a high flare load in the case of a fire scenario leads to a high back pressure of the PSV, causing severe damage to the PSV. Thus, the study indicated that the diameter of the tail pipe must be increased or the tail pipe should be divided into two separate pipes to reduce the flare load for each pipe. Unfortunately, this study did not
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consider limit of the Mach number that affects the vibration of the pipeline. In another study, Sisshartha Mukherjee (Mukherjee, 2008) numerically optimized flare
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networks in the case of cooling-water failure and external fire scenarios by solving governing
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equations. As one of the constraints for optimization, the maximum values of the PSV back pressure and Mach number in the header were considered. For estimation of these parameters,
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the American Petroleum Institute (API) Standard 521 was employed. During optimization, design variables, including the number of PSVs, configuration of the pipelines, and diameters of
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the pipes, were adjusted, satisfying the proposed constraints. However, only a steady-state model was developed; dynamic behavior, such as changes in the relieving loads over time, was not
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considered. Thus, the design margin was too large for designing pipes with the optimum size for flare networks.
Recently, Prashanth Siddhamshetty (Siddhamshetty et al., 2019) performed a simplified model
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of ignition phenomenon at a wellhead blowout situation. In this study, they reviewed previous papers comprehensively to analyze the relationship between a set of parameters and burning efficiency. Moreover, they developed a simplified model to calculate the burning efficiency, and the results showed high accuracy. After PSV opening, the liquid is generally formed in the pipe due to the depressurization. This liquid formation affects the flame length at the flare stack and
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flame temperature. In terms of the safety, prediction of flame length and the flame temperature is considered to be quite important. Most flare systems of existing chemical processes were designed according to the steady-state condition. This method uses the sum of flare loads, which are collected in header pipes from different sources of PSVs, to design the diameters of the headers, as described in Figure 2(a).
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However, the flare load that is released from each PSV has a peak point with regard to the flowrate immediately after its release, and the flowrate decreases, as shown in Figure 2(b). Thus,
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if the diameters of the headers are designed according to the estimated maximum flowrates, the
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design margin becomes too large. Additionally, all the main equipment is installed with long distances for safety, and it takes different times for the released fluid from each PSV that is
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installed in the equipment to reach the main header. Consequently, the actual cumulative flare load that passes through the main header is smaller than algebraic sum of the flare loads from the
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PSVs. Therefore, it is necessary to estimate the actual flare loads that pass through the main headers over time from the point of their release until they reach the flare tips. However, few
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studies involving the analysis of flare systems with dynamic simulation have been published.
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(b) Figure 2. Calculation of the total flare load using the (a) steady-state model and (b) dynamic
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model
Gilardi and Tonello (Gilardi and Tonello, 2011) estimated the total flare load along the main
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flare header over time in a power-failure scenario by using a dynamic simulation model. The results indicated that the maximum accumulated flare load that was collected from various
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sources of PSVs was significantly smaller than the algebraic sum of the flare loads from the individual PSV sources. Alban Sirven et al. (Sirven et al., 2011) estimated the total flare load in the case of reflux failure of a crude distillation unit using a DynaSim simulation. The size and number of PSVs were optimized to minimize the total capital cost of the flare system. However, the study focused on the total flare load but did not consider the Mach number of the fluid in the
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(a)
pipes. Ali Shafaghat (Shafaghat, 2016) performed a dynamic simulation of the flare system in an external-fire scenario and analyzed the changes in the fluid properties over time with regard to the pressure, temperature, and mass flowrate during the depressurization period. Then, the estimated Mach numbers from steady-state and dynamic simulation models were compared. The
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estimated Mach number from the steady-state model exceeded the design criteria, whereas the Mach number from the dynamic model was significantly smaller than the design criteria during the whole relieving period, except for the initial 1 min. Thus, the size of the header pipes could be reduced. However, only the main header was considered in the study. Most previous studies on the design of the flare network system involving dynamic simulation
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have mainly focused on comparing the flare loads, Mach numbers, back pressures, and sizes of the main header pipes in particular emergency scenarios. They seldom revealed the optimum size
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of the whole flare system, including the tail pipes that are connected to PSVs. Furthermore,
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onshore chemical plants were mainly considered by researchers, while its impact on capital cost
2.
Modeling methods
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2.1. General PSV relieving scenarios
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saving is far larger for offshore plants.
In general, a PSV opens owing to overpressure of the equipment, which has various causes.
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Therefore, it is necessary to consider each scenario that causes overpressure of the equipment and apply the results to the design of the flare network system to ensure its safe operation. Relieving scenarios can be divided into two main categories. The first is the continuous-emission
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scenario, which includes purging, control valve leakage, and process venting, and the second is the emergency scenario, in which the flare load is released in a short time for depressurization, e.g., external fires, power failure, and outlet blockages (Davoudi et al., 2014). In general, the emergency scenario is employed to determine the size of the flare network pipes, because the flare relieving load for the continuous-emission scenario is significantly smaller than that for the emergency scenario (Sangsaraki and Anajafi, 2015). The details and design guidelines of the
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typical emergency scenarios, such as closed outlets, cooling-water failure in the condenser, overfilling, failure of automatic controls, fire, and power failure, are provided in API Standard 521. 2.2. Steady-state modeling A typical flare network system can be modeled in the steady state by using the AFSA. The
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AFSA, which is specialized for the design of the flare network system, can be used to estimate the Mach number and momentum of the fluid in pipes and the noise of the pipes. Furthermore, a
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design engineer can easily verify that the design criteria are satisfied as he/she adjusts the inputs
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of the flare network system. In the AFSA, an engineer must specify inputs, such as the expected relieving load of each PSV, set pressures of the PSVs, orifice sizes of the PSVs, and discharge
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pressures of the PSVs. Once all inputs are entered in the simulation model, the design
estimated. 2.3. Dynamic modeling
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parameters, such as the Mach number, back pressure, and noise level of the pipes, can be
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In contrast to the traditional steady-state model of the flare network system, a dynamic model can analyze the changes in the flare loads and their impact on the back pressure of the PSV and Mach number of the fluid in the pipes over time. Therefore, the estimated results are regarded as
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more accurate and reliable than those from the steady-state model for designing the flare network system. Additionally, it is possible to accurately estimate the uneven distribution of flare loads between PSVs that are installed on the same equipment and to observe the changes in the pressure, temperature, and flowrates under emergency conditions. One of the most important variables in designing the flare network system is the pressure change of the equipment, and the opening tendency of PSVs can be estimated over time according to the pressure change inside a
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vessel. The factors that affect the pressure of the vessel are the initial operating conditions of the feed stream that enters the vessel (i.e., flowrate, temperature, pressure, and composition), size of the vessel, and closing or opening speed of the valves that are installed in the inlet and outlet pipes of the vessel. In this study, gPROMS Process Builder was used in order to simulate the dynamic state. The gPROMS is based on the dynamic state, and calculate using Equation
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Oriented (EO) method. Therefore, both the flowsheet based model and the custom model are simulated simultaneously. As a result, the calculation speed is much faster than Sequential
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Modulate (SM) method.
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2.4. Estimation of Mach number
The Mach number is defined as the ratio of the velocity of a fluid to the velocity of sound in
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the fluid. In the steady-state model (i.e., AFSA), the Mach number is calculated using Equation
Equation 2.
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1, and the sonic velocity is calculated according to the pressure and density differences using
𝐹𝑙𝑢𝑖𝑑 𝑣𝑒𝑙𝑜𝑐𝑖𝑡𝑦 (1) 𝑆𝑜𝑛𝑖𝑐 𝑣𝑒𝑙𝑜𝑐𝑖𝑡𝑦 𝑖𝑛 𝑡ℎ𝑒 𝑓𝑙𝑢𝑖𝑑 ∆𝑃 (2) 𝑐2 = ∆𝜌 Here, Ma represents the Mach number, c represents the sonic velocity in the fluid, and ΔP and
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𝑀𝑎 =
Δρ represent the differences in the pressure and density, respectively, between the pipe inlet and
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outlet.
In the dynamic model (i.e., gPROMS ProcessBuilder), the Mach number was calculated using
Equation 3, which is described in the API Standard 521. Equation 1 can also be used to calculate the Mach number in the dynamic model; however, API Standard 521 is more widely used by engineers in industry.
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𝑚̇ 𝑍∙𝑇 (3) 𝑀𝑎 = 3.23 × 10−5 ( )( ) 2 𝑃∙𝐷 𝑀𝑤 Here, 𝑚̇ represents the mass flowrate in the pipe, P represents the pressure of the pipe outlet, D represents the diameter of the pipe, Z is the compressibility factor, T represents the temperature of the pipe outlet, and Mw represents the average molecular weight.
Scenario analysis
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3.1. Separator outlet blocked discharge
Figure 3. Schematic of the "HP separator outlet blocked discharge” scenario In this study, “high-pressure separator (HP separator) outlet blocked discharge” was selected
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as a governing scenario. The feed stream to an HP separator is directly connected to a riser from the subsea region, and operating pressure of the fluid is very high. Therefore, if the outlet valve of the separator is blocked, the risk of explosion becomes high owing to the high-pressure stream inflow to the equipment. Figure 3 presents a schematic of the “HP separator outlet blocked discharge” scenario. The hydrocarbon mixture passes through the HP separator, and its pressure
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and temperature are approximately 14.8 bar and 70 ºC, respectively. If the outlet valves are blocked at a certain point, the fluid that continuously enters the vessel accumulates inside the separator, increasing the pressure inside the vessel. Then, as the separator pressure reaches the pressure alarm high high (PAHH) point, the emergency shutdown valve (ESDV), which is connected to the inlet stream, is scheduled to close for approximately 20 s to block the inlet
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stream. While the ESDV closes, the inlet flow continuously enters the HP separator. Finally, the HP separator pressure reaches the set pressure of the PSVs, and the PSVs opens to release the
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fluid inside the vessel and reduce its pressure. The released fluid initially passes through the tail
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pipes of the PSVs, and it gathers in the main flare header pipes. The fluid transported through the
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main header pipes is discharged into the atmosphere after incineration at the flare stack.
Table 1 presents the maximum allowable working pressure (MAWP) and PAHH of the HP
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separator, and Table 2 presents the set pressure of the PSVs installed on the top of the vessel. The sizes of the PSV orifices are identical, whereas the set pressure differs among the PSVs.
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According to API 521, the maximum allowable set pressure of the 1st PSV should be equal to the MAWP, and those of the supplemental PSVs should be <105% of the MAWP. Therefore, the
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highest set pressure was set as 104.5% of the MAWP in this case.
Table 1. Design of the HP separator Item
Pressure [bar]
MAWP
26.634
PAHH
21.690
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Table 2. Set pressure of each PSV Set pressure [bar]
PSV-A
26.634
PSV-B
27.034
PSV-C
27.434
PSV-D
27.834
PSV-E (spare)
26.634
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Item
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3.2. Development of steady-state model
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Figure 4 shows a snapshot of the steady-state model developed using the AFSA for the “HP separator with outlet blocked discharge” scenario. The configuration of the model in Figure 4
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includes PSVs, tail pipes, main header pipes, a flare knockout drum, and a flare stack. The input variables at each PSV, such as the set pressure, orifice size, and mass flowrate of flare load, were
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directly specified in the software. For the modeling of pipes, the design parameters, such as the
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diameter, thickness, materials, and length of each pipe segment, must be specified by a user.
Figure 4. Configuration of the "HP separator outlet blocked discharge" scenario in the AFSA In this study, modulating pilot-operated PSVs identical to those in the actual industrial project were selected. The pilot-operated PSVs mainly consist of a pilot piston and a main valve. The pilot piston presses the main valve through the piston using the vessel pressure. The top area of
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the pilot piston is larger than the bottom area that is in contact with the main valve. Therefore, the net force presses the pilot piston tightly until the vessel pressure reaches the set pressure. When the vessel pressure reaches the set pressure, the pressure of the pilot piston is vented for reduction of the seating force, and the main valve piston is lifted for depressurization of the vessel. Because the pilot-operated PSV is operated using the vessel pressure, the lift of the main
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valve is not significantly affected by the back pressure (American Petroleum Industry, 2008; Scott and MacKinnin, 2011). For estimation of the relieving load from each PSV, a mass
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flowrate of 267,200 kg/h was specified for each PSV model. For this, the total feed flowrate to
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the HP separator, as shown in table 3, was divided by the number of PSVs, in accordance with the API Standard 521. Two different sizes of tail pipes (10 inches for the 1st tail pipe and 12
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inches for the 2nd tail pipe) were specified in the tail pipe model. A diameter of 24 inches was used for the header pipes, and a diameter of 30 inches was used for the main header pipe
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immediately before the knockout drum. The composition of the relieving fluid (mole fractions) was obtained from the industry project data, as shown in Table 4. The feed condition such as
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flowrate, temperature and pressure of each PSV was assumed by a process engineer, and they were specified as an input for each PSV in the flowsheet. Therefore, the specified feed condition
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of each PSV was referred as relieving condition in a steady-state model.
Table 3. Steady state conditions Relieving conditions Pressure (bar)
26.634
Temperature (ºC)
70
Mass flowrate (kg/h)
1,336,000
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Table 4. Mole fractions under the relieving conditions Mole fraction
Component
Mole fraction
Component
Mole fraction
CO2
0.0029
C6*
0.0041
C15*
0.0032
Nitrogen
0.0019
C7*
0.0064
C16*
0.0026
Methane
0.3490
C8*
0.0062
C17*
0.0023
Ethane
0.0267
C9*
0.0049
C18*
Propane
0.0138
C10*
0.0044
C19*
i-Butane
0.0032
C11*
0.0037
n-Butane
0.0064
C12*
0.0034
i-Pentane
0.0028
C13*
0.0036
n-Pentane
0.0030
C14*
0.0030
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Component
0.0023 0.0022 0.0225
H2S
0.0001
H2O
0.5153
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C20+*
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3.3. Development of dynamic model
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Figure 5. HP separator blocked discharge scenario model in gPROMS
As a next step, the dynamic model was developed using gPROMS ProcessBuilder for the “HP
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separator with outlet blocked discharge” scenario. In contrast to the steady-state model that was developed previously, the main process was included in the model together with the flare network system, as shown in Figure 5. In particular, the dynamic model included a riser stream from the reservoir to the HP separator, the size of the main equipment, and the pressure drops
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through the valves and pipes. Tables 5 and 6 present the operating conditions of the feed stream and its composition (mole fractions), respectively. In contrast to steady-state model, feed condition to PSV is not fixed in dynamic simulation model. In this case, the topside process is integrated with subsea piping system from the reservoir. For the reason, reliving condition of the PSV is estimated through simulation based on the operating condition when an emergency situation occurs. For dynamic simulation of the scenario, an “outlet valve” block was
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implemented in the model, and it was manually closed once the normal operation was stabilized. The system was stabilized 50 s after the dynamic simulation started. When the pressure of the separator reached the PAHH point (21.690 bar) from its initial point (14.8 bar) owing to the blockage of the outlet valve, the ESDV closed for 20 s. Then, four PSVs, which were installed on the HP separator (excluding the spare PSV), started to open in a row once the pressure of the
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HP separator reached the set pressure of each PSV, and the compressed fluid was released into the tail pipes. As expected, the relieving load changed significantly over time after the PSV
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opened, until it was closed again. Furthermore, the rupture of one PSV affected the other PSVs
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that were connected via the pipes, owing to the interaction of the back pressure and packing
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effect.
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Table 5. Feed stream conditions of dynamic simulation Operating conditions
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Temperature (ºC)
70
Mass flowrate (kg/h)
900,000
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Pressure (bar)
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Table 6. Mole fractions of the feed stream Mole fraction
Component
Mole fraction
Component
Mole fraction
CO2
0.0043
C6*
0.0035
C15*
0.0022
Nitrogen
0.0029
C7*
0.0055
C16*
0.0020
Methane
0.3190
C8*
0.0052
C17*
0.0019
Ethane
0.0273
C9*
0.0031
C18*
0.0013
Propane
0.0137
C10*
0.0036
C19*
i-Butane
0.0027
C11*
0.0028
C20+*
n-Butane
0.0059
C12*
0.0035
i-Pentane
0.0027
C13*
0.0030
n-Pentane
0.0031
C14*
0.0027
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0.0020 0.0022 0.0002
H2O
0.5735
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H2S
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Results and discussion
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The simulation results of the dynamic model were analyzed. First, the dynamic behavior of the HP separator was analyzed. Then, the dynamic behavior of the flare system, such as the opening
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of the PSVs and the relieving loads from the PSVs, was analyzed. Lastly, the results of the dynamic simulation were compared with those of the steady-state model, and the optimum sizes of the PSV orifices and pipes were estimated according to the sizing criteria of API Standard 521. In this study, the following CPU and GPU were used for computation. CPU: Intel® Xeon® Gold 5120 CPU 2.20GHz
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GPU: NVIDIA GeForce GTX 1080 Ti
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In general, 3 – 5 seconds were taken to perform simulation of the steady-state model, while
200 – 210 seconds for dynamic model. Because the dynamic model considered both the main process and flare network simultaneously, it requires more time for computation. However, the
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simulation results of the dynamic model showed much higher accuracy compared to the ones from the steady-state model. 4.1. Dynamic behavior of HP separator As shown in Figure 6(a), the HP separator normally operated at a constant pressure of 14.8 bar. Once the gas and liquid outlet valves of the HP separator were blocked at 50 s, as shown in
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Figure 6(b), the pressure of the vessel increased steadily to 27.6 bar (at 80 s) owing to the continuous entrance of the feed stream to the HP separator. For reference, it took approximately
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10 s for the outlet valves to close fully. When the pressure of the HP separator reached the
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PAHH point (21.690 bar) at approximately 67 s, the ESDV closed for approximately 20 s, and the inlet mass flowrate started to decrease rapidly. Because of the continuous inflow of the feed
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stream during the closing period of the ESDV, the pressure continuously increased and reached the PSV-A set pressure (26.634 bar) at 76 s. PSV-A opened rapidly, and a mixture of gas and
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liquids was released to protect the HP separator from further build-up of pressure. Although PSV-A opened, the pressure was accumulated because the opening of only one PSV was
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insufficient to prevent further pressure build-up. Therefore, the pressure slowly increased further and ultimately reached the set pressures of PSV-B (27.034 bar) and PSV-C (27.434 bar). Thus, PSV-B and PSV-C sequentially opened at 77 and 78 s, respectively. On the other hand, PSV-D
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did not open, because the vessel pressure started to decrease afterwards. After the pressure decreased to 26.60 bar, it remained constant below the MAWP (26.634 bar) from 105 s because the SDV was fully closed at 87 s. According to API Standard 521, the maximum allowable overpressure in the case of multiple valve installation should be <116% of the MAWP. Because the highest pressure of the vessel was 27.6 bar, the maximum overpressure was 103.6%, which satisfied the standard.
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(b)
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(a)
Figure 6. Dynamic behavior of the HP separator: (a) mass flowrate and pressure; (b) opening of
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block valve and ESDV
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4.2. Dynamic behavior of PSVs
Figure 7(a) illustrates the dynamic behavior of the PSVs installed on the HP separator. The
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simulation results indicated that three of the four PSVs (with the exception of the spare) opened
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to protect the vessel. The specified set pressures of PSV-A, PSV-B, PSV-C, and PSV-D were 26.634, 27.034, 27.434, and 27.834 bar, respectively, and the valves opened sequentially over
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time. PSV-A opened at 76 s, PSV-B opened at 77 s, and PSV-C opened at 78 s, and the opened PSVs closed in the reverse order (i.e., C, B, and A). PSV-A released the largest flare load (27.96 kg/s), and PSV-B and PSV-C released smaller flare loads (16.59 and 5.54 kg/s, respectively).
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As described previously, the lift of the pilot-operated PSV was not affected by the back pressure. However, the relieving load of the PSV and the Mach number of the fluid in the tail or header pipes were affected by the back pressure. The back pressure of the PSVs was calculated using the operating pressure of the flare knockout drum, which normally operated at 1 atm, by adding pressure drops through the pipes (Zadakbar et al., 2015; Hernández-Suárez et al., 2007). Figure 7(b) shows the back pressures of PSV-A, B, and C. As PSV-A opened, the back pressures
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of PSV-B and PSV-C were affected simultaneously. The differences in the flare loads of the
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PSVs and the back pressures caused the differences in the Mach number of the fluid in the pipes.
(b)
(a)
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Figure 7 PSV conditions: (a) mass flowrate; (b) back pressure
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4.3. Mach-number calculation results and comparison with AFSA
A profile of the estimated Mach numbers and the released flare load and outlet pressure of the
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PSV at one of the tail pipes are shown in Figures 8(a) and 8(b), respectively. Once the PSV opened, the Mach number increased rapidly and decreased after reaching the peak point. To
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obtain a conservative design for the sizing of the flare pipes, the maximum values of the Mach
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numbers were considered in this study.
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(b)
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(a)
Figure 8. (a) Estimated Mach number and (b) released flare load from the PSV and its outlet
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pressure
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In general, the estimated Mach numbers differed between the inlet and outlet points of the same pipe. Because the mass flowrate along the pipe was constant and the pressure was lower at
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the outlet point owing to the pressure drop, the estimated Mach number at the inlet was larger
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the pipe was mainly considered.
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than that at the outlet, as indicated by Equation 3. In this study, the Mach number at the outlet of
The estimated Mach numbers from the steady-state and dynamic models are compared in Figure 9 and Table 7. Overall, the Mach numbers of the dynamic model were smaller than those of the steady-state model, except for pipes 1–4, which were connected to PSV-A and PSV-B.
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The estimated Mach numbers of pipes 1–4 in the dynamic model were larger than those in the steady-state model. Moreover, pipes 1 and 2, which were connected to PSV-A, exceeded the line sizing criteria of 0.7, which can cause safety problems, such as the vibration of the pipe during operation. This is because the PSVs had the same orifice size but different set pressures were specified, which resulted in the release of different amounts of flare loads from the PSVs. Because the largest flare load was released from PSV-A, which had the lowest set pressure, the
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largest Mach number was estimated in the tail pipes connected to PSV-A. Additionally, the Mach numbers of PSV-B and C, which opened later owing to the high set pressure, were estimated to be lower than that of PSV-A. The Mach number in the tail pipes from PSV-D was
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estimated to be zero because PSV-D was not open, owing to the relatively high set pressure.
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Figure 9. Comparison of the Mach number between the steady-state and dynamic models Table 7. Calculated Mach number of each pipe Tail pipe
PSV-B
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PSV-A
Header
PSV-C
PSV-D
pipe
pipe
pipe
pipe
pipe
pipe
pipe
pipe
pipe
pipe
pipe
pipe
pipe
1
2
3
4
5
6
7
8
12
13
14
15
16
Steadystate
0.6
0.4
0.57
0.37
0.56
0.37
0.57
0.37
0.09
0.19
0.29
0.44
0.27
Dynamic
0.87
0.80
0.74
0.59
0.30
0.21
0.00
0.00
0.00
0.05
0.19
0.47
0.32
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4.4. Pipe-size optimization results The optimum sizes of the tail and main header pipes were estimated according to the dynamic simulation results and the line sizing criteria of API Standard 521. In practice, the sizes of tail pipes and sub-header pipes should be adjusted to satisfy the maximum value of the Mach number (0.7). For the sizing of the main header, the estimated Mach number should be <0.5. Detailed
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guidelines for the line sizing are presented below (Kerr et al., 2018). Line sizing criteria in API 521: Mach number of header < 0.5
•
Mach number of sub-header and tail pipe < 0.7
•
MABP: 50% of relief-valve set pressure for modulating pilot-operated PSV
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•
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In this study, Equation 3 was used to determine the appropriate size of the pipes by adopting
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the maximum Mach numbers (0.5 or 0.7). The results are shown in Figure 10 and Table 8.
Figure 10. Calculated pipe diameter vs. original design
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Table 8. Calculated diameter of each pipe Tail pipe Header PSV-A
PSV-B
PSV-C
PSV-D
pipe 2
pipe 3
pipe 4
pipe 5
pipe 6
pipe 7
pipe 8
pipe 12
pipe 13
pipe 14
pipe 15
pipe 16
Original design
10
12
10
12
10
12
10
12
24
24
24
24
30
Dynamic
14
14
10
10
5
6
0
0
0
8
14
24
24
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pipe 1
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As shown in Figure 10, there were many chances to reduce size of the pipes compared with the
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original design. The diameter of most pipes could be reduced or maintained. In the case of tail pipes 1 and 2, which were connected to PSV-A, the diameter must be increased by one or two
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sizes to ensure safe operation. As described previously, the dynamic simulation results proved that larger flare loads were expected to be released through PSV-A than through the other PSVs,
Conclusion
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5.
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whereas in the steady-state model, the same flare load was allocated to each PSV.
A governing scenario of “HP separator outlet blocked discharge” in an offshore plant was adopted to compare the design results of a flare network system between steady-state and dynamic models. For the steady-state model, the widely used software of the AFSA was used,
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whereas gPROMS ProcessBuilder was used to develop the dynamic simulation model. Compared with the steady-state model, the dynamic model estimated the behavior of the relieved fluids over time more accurately, including the flowrates, pressures, and temperatures of the fluids in the pipes. Detailed information regarding the main process, process control scheme, and ratings of major equipment were incorporated into flare network system. Then, the Mach
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numbers of the fluids in the pipes were estimated for both the dynamic and steady-state simulations, and the results were compared. The results indicated that the original design of the flare network must be modified significantly for capital cost saving and safe operation. Lastly, the optimum sizes of the tail pipes, sub-headers, and main headers were calculated according to both the dynamic simulation results and the standard line sizing criteria from API Standard 521.
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The results indicated that the diameters of most pipes could be reduced, reducing the capital cost. However, the diameters of a few pipes must be increased to avoid vibration and noise of the
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pipes during abnormal operation.
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For sizing of the main header and sub-header pipes, the most extreme case of the emergency situations should be considered as a governing scenario. Because the “HP separator outlet
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blocked discharge” was considered to be the governing case among all potential emergency situations (e.g., power failure, control failure, fire, etc.) in this study, the size of the main header
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and sub-headers should be big enough to to cover all other emergency situations. Furthermore, performance of the control system was examined through dynamic simulation to avoid any
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potential failure of the system.
The refinery, petrochemical and off-shore plants have flare network system to protect the units and humans in abnormal operations. Generally, the flare network system in these plants is
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designed based on the steady-state model. Therefore, high margin is added to the design of flare network system to cover inaccuracy that is caused by the steady-state model. However, as computation speed increases rapidly these days, computational time for dynamic simulation becomes much shortened. For the reason, change of operating condition can be examined more accurately during emergency situation by using dynamic simulation, and we can design more robust and safer flare network system with lower capital cost. This proposed methodology allows
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us to design the system efficiently with high accuracy. However, it requires substantial information regarding the process and the ratings of the equipment, as well as considerable manhours. Therefore, a new methodology should be developed for designing the flare network system more efficiency while maintaining high accuracy with reduced man-hours.
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Funding This research was supported by a Korea Institute for Advancement of Technology (KIAT) grant
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funded by the Korean Government (MOTIE) (P0008475, Development Program for Smart
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Digital Engineering Specialist). It was also supported by the Basic Science Research Program through the National Research Foundation of Korea (NRF) funded by the Ministry of Education
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(2019R1F1A1058979).
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Declaration of interests
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The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.
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REFERENCES American Petroleum Industry., 2008. Sizing , Selection , and Installation of Pressure-Relieving Devices in Refineries Part I — Sizing and Selection, eight ed. Washington, DC. American Petroleum Industry., 2014. Pressure-relieving and Depressuring system, six ed.
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Washington, DC.
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Bader, A., Baukal, C. E., Bussman, W., 2011. Selecting the proper flare systems. Chem. Eng. Prog., 107(7), 45–50.
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Bjeere, M., Eriksen, J. G. I. ., Andreaden, A., Stegelmann, C., Mando, M., 2016. Analysis of Pressure Safety Valves for fire protection on offshore oil and gas installations. Process. Saf.
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Environ. Prot., 105(Iso 23251), 60–68. doi:10.1016/j.psep.2016.10.008 Chopinet, F., Marcec-Rahelic, N., 2015. Upgrade of a Flare Network for a New Refinery Unit.
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Goriva I Maziva, 54(1), 32–43.
Davoudi, M., Aleghafouri, A., Safadoost, A., 2014. Flaring networks assessment in South Pars
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Gas processing plant. J. Nat. Gas. Sci. Eng., 21, 221–229. doi:10.1016/j.jngse.2014.08.008 Gilardi, C., Tonello, M., 2011. Dynamic simulation, HIPS & combined probability analysis reduce flare loads. Associazione Termotecnica Italiana(ATI), Milano, Italy, March 31, 2011.
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Goyal, R. K., Al-Ansari, E. G., 2009. Impact of emergency shutdown devices on relief system sizing and design. J. Loss. Prev. Process. Ind., 22(1), 35–44. doi:10.1016/j.jlp.2008.07.008
Hernández-Suárez, R., Puebla, H., Aguilar-López, R., 2007. Parametric approach for the optimal design of knockout drums. Ind. Eng. Chem. Res., 46(21), 7008–7017. doi:10.1021/ie0701930
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Jin, X., Ray, A., Edwards, R. M., 2010, Integrated Robust and Resilient Control of Nuclear Power Plants for Operational Safety and High Performance, IEEE Trans. Nucl. Sci., 57(2), 807-817. doi:10.1109/TNS.2010.2042071 Khan, F., Rathnayaka, S., Ahmed, S., 2015, Methods and models in process safety and risk management: Past, present and future, Process Saf. Environ., 98, 116-147. doi: 10.1016/j.psep.2015.07.005
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Kerr, T. A., Myers, D. B., Piggott, B. D., Vance, A. E., 2018. Solving pressure relief valve and pipng capacity problems. 2018 GPA Midstream Convection, San Antonio, U.S.A. April 15-
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Lee, H. S., Ko, B. S., Yang, J. M., Lee, C. J., Yoo, J. H., Shin, D. G., Park, C. H., Ko, J. W., 2012. Reduction of thermal radiation by steam in flare stack system. Korean. J. Chem. Eng.,
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Mukherjee, S., 2008. Pressure-Relief System Design. Online, Chem. Eng., 40-45.
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https://www.chemengonline.com/pressure-relief-system-design/ (accessed August 18, 2019) Roberts, T. A., Buckland, I., Shirvill, L. C., Lowesmith, B. J., Salater, P., 2004. Design and Protection of Pressure Systems to Withstand Severe Fires. Process. Saf. Environ. Prot.,
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82(2), 89–96. doi:10.1205/095758204322972735 Sahoo, M., 2013. High Back Pressure on Pressure Safety Valves (PSVs) in a Flare System. University of Bergen, Norway, Master Thesis.
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Sangsaraki, M. E., Anajafi, E., 2015. Design Criteria and Simulation of Flare Gas Recovery System. International Conference on Chemical, Food and Enviroment Engineering. Dubai, January 11-12, 2015 doi:10.17758/iaast.a0115008
Scott, J. R., MacKinnin, N., 2011. Pilot-operated safety relief valves: A simple, effective plant upgrade. Online, Hydrocarb. Process., 1–7. https://www.hydrocarbonprocessing.com/magazine/2011/november-2011/special-report-
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plant-safety-and-environment/pilot-operated-safety-relief-valves-a-simple-effective-plantupgrade (accessed August 27, 2019) Shafaghat, A., 2016. Depressurization Dynamic Modeling and Effect on Flare Flame Distortion. University of Calgary, Canada, Master Thesis. Shin, C. H., Yoon, Y., Park, J. H., Ma, B. C., 2018. Design of the safety standard at hydrofluoric acid handling facilities for risk reduction. Korean. J. Chem. Eng., 35(5), 1225–1230.
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Siddhamshetty, P., Ahammad, M., Hasan, R., Kwon, J., 2019. Understanding wellhead ignition as a blowout response. Fuel., 243, 622-629. doi: 10.1016/j.fuel.2019.01.142
Understanding. Online, DIGIRAL Refin.
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Sirven, A., Grosclaude, J., Fenol, G., Saada, J., 2011. Optimising safety relief and flare systems
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https://www.digitalrefining.com/article/1000394#.XVkKrugzaCq (accessed August 27, 2019)
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Song, X., Cui, L., Cao, M., Cao, W., Park, Y., Dempster, W. M., 2014. A CFD analysis of the dynamics of a direct-operated safety relief valve mounted on a pressure vessel. Energy.
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Convers. Manag., 81, 407–419. doi:10.1016/j.enconman.2014.02.021 Zadakbar, O., Khan, F., Imtiaz, S., 2015. Development of Economic Consequence Methodology for Process Risk Analysis. Risk. Anal., 35(4), 713–731. doi:10.1111/risa.12313 Zhu, P., Liyanage, J. P., Panesar, S. S., Kumar, R., 2019. Review of workflows of emergency
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PII:
S0957-5820(19)31781-1
DOI:
https://doi.org/10.1016/j.psep.2019.12.008
Reference:
PSEP 2022
To appear in:
Process Safety and Environmental Protection
Received Date:
8 September 2019
Revised Date:
18 November 2019
Accepted Date:
9 December 2019
Please cite this article as: Jo Y-pyeong, Cho Y, Hwang S, Dynamic analysis and optimization of flare network system for topside process of offshore plant, Process Safety and Environmental Protection (2019), doi: https://doi.org/10.1016/j.psep.2019.12.008
This is a PDF file of an article that has undergone enhancements after acceptance, such as the addition of a cover page and metadata, and formatting for readability, but it is not yet the definitive version of record. This version will undergo additional copyediting, typesetting and review before it is published in its final form, but we are providing this version to give early visibility of the article. Please note that, during the production process, errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain. © 2019 Published by Elsevier.
Dynamic analysis and optimization of flare network
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Yeon-pyeong Jo, Yongheon Cho and Sungwon Hwang*
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system for topside process of offshore plant
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Department of Chemistry and Chemical Engineering, Inha University, Incheon 22212, South
Corresponding Author
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Korea
Highlights
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* E-mail address: [email protected].
Design and optimization of flare relief system in emergency situation at offshore plant
Designing flare relief system at steady-state has an excessive margin for line size
Reduction of overall capital cost is achieved by using dynamic simulation
Enhancement of flare system safety considering uneven distribution of flare loads
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ABSTRACT
This study introduces a new approach for designing a flare network system that ensures economic feasibility and safety using dynamic simulation analysis with the gPROMS ProcessBuilder software. A case of “separator outlet blocked discharge” was selected as a
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pressure-relief scenario based on the American Petroleum Institute Standard 521, and design data from a previous offshore project were used. As the main process and flare network system were
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dynamically simulated, the effects of the pressure, temperature, and liquid-level changes in the vessel on the opening of the pressure safety valve (PSV) were analyzed. Then, dynamic
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simulation results, including those for the flare load, PSV back pressure, and Mach number, were
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compared with those obtained from a steady-state model employing the Aspen Flare System Analyzer. Lastly, the sizes of the PSVs, branch lines, and main headers were optimized to
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minimize the overall capital costs and ensure the safety of the flare network system. This methodology can be applied to all existing and newly designed flare network system to enhance
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safety and reduce capital costs.
Keywords: Emergency flare emission, Dynamic simulation, Mach number calculation, Line
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sizing criteria, Pipe line optimization
1.
Introduction
In large-scale chemical plants, various types of equipment operate, mainly including rotating
machines, separators, reactors, heat exchangers, and vessels. They normally run under the allowable operating conditions (e.g., temperatures, pressures, and flowrates) specified by either the manufactures or the operators. However, in emergency situations, such as fires, utility
2
failures, and outlet blockages, such equipment tends to operate abnormally, leading to overpressure of the equipment (American Petroleum Industry, 2014). This may result in serious fires and explosions, causing severe damage to humans, property, and the environment (Shin et al., 2018). To prevent such accidents, pressure-relief systems are employed in most chemical plants. This pressure-relief systems protect the process equipment and relieve overpressure that
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arises during normal or abnormal operation. Furthermore, the emergency shutdown (ESD) system is widely used to protect the equipment
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and humans. Pengyou Zhu (Zhu et al., 2019) developed an overall framework to identify gaps
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and challenges of ESD system based on the field study in Norway. The authors mentioned that the workflow of ESD system includes three key parameters (i.e., technical, operational,
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organizational parameters). The technical parameter is related to equipment and systems, the organizational parameter to personnel with defined roles or functions and specific competence,
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and the operational parameter to actions and activities that the personnel has to perform. According to identification results using the framework, the authors found 15 challenges during
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the operation and maintenance of ESD system. Among the 15 challenges, three, seven and five of them were related to technical, organizational, and operational challenges, respectively. As a result, the paper provided effective guideline regarding the operation and maintenance of ESD
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system.
In addition, the emergency situation can be handled by applying an integrated robust and
resilient control scheme to the process. In such a case, the role of robust control is to guarantee the product specifications in emergency situation, and the role of resilient control is to stabilize the process fast during an emergency situation (Jin et al, 2010). However, further research is
3
necessary, and it is important that the flare network system should be integrated with such control system. Moreover, in terms of risk assessment, the emergency scenario can be predicted based on historical data. Faisal Khan (Khan et al., 2015) presented the overall risk assessment method. In the study, the authors adopted various methods to manage the risk of the process. One of the
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methods, which used the historical data, was to combine the Fault Tree Analysis (FTA) with the Event Tree Analysis (ETA). FTA graphically illustrated the logical relationship between cause,
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consequence and failure paths. ETA evaluated the sequence of failure processes and analyzed its
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consequences. The authors described System Hazard Identification, Prediction and Prevention (SHIPP), which combined FTA and ETA. SHIPP is able to predict the probability of emergency
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by analyzing the history of emergency. As a results, the frequency of emergency situations can be predicted.
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Figure 1 shows the simplified structure of the pressure-relief system. It contains pressure-relief devices, headers and tail pipes that transfer flare loads, knock-out drums that separate liquid and
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vapor before combustion, and a flare stack that burns flare gas safely (Bader et al., 2011). Pressure-relief devices typically include a pressure safety valve (PSV), a blowdown valve (BDV), and a rupture disk, and they release fluid from the inside of the equipment to relieve
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overpressure. The PSV performs depressurization automatically under abnormal conditions that cannot be recognized by the operator. It normally opens via spring or elastic action when the pressure of the equipment reaches a designated set pressure. The BDV also performs depressurization, but it is generally used for manual operation to release fluid quickly from the inside of the vessel for process shutdown or equipment replacement. While the PSV constantly maintains the pressure of the vessel below the set pressure, the BDV reduces the pressure of the
4
vessel to the level set by the operator (Bjeere et al, 2016). A rupture disc is installed in facilities or devices, such as pressure containers, pipes, and reactors, and it is ruptured when the pressure reaches the set point during operation to protect the equipment. Thus, it is only used once and should be replaced with a new one once it is ruptured. In general, it is used as a spare device of the PSV when the valve function is impaired because abnormal materials have accumulated in
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the PSV and the valve cannot respond quickly enough. The most commonly used relief device in emergency situations is the PSV (Roberts et al, 2004). The design of the PSV depends on the
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size of the orifice, the flare load, and its back pressure. Therefore, it is important to install PSVs
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of the correct type and size, as well as the correct number of PSVs, for safe depressurization.
Figure 1. Typical structure of the flare relief system
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Flare gas released from each pressure-relief device, which is installed in various equipment, flows through a tail pipe and a sub-header and gathers in the main header. Subsequently, the relieved vapor and liquid are collected in a knockout drum (K.O Drum). Then, gas is eventually sent to the flare stack for incineration before its release to the atmosphere (Lee et al., 2012). During normal operation of the flare system, when there is no released fluid from the PSV, small
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amounts of light hydrocarbon vapor flow through the header to maintain the operation of the flare network (Chopinet and Marcec-Rahelic, 2015). However, in the case of an emergency or
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abnormal operation, large amounts of released fluids flow through the header with a high
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velocity at a high pressure (Song et al., 2014). In general, the following aspects should be
Limits of the PSV back pressure
•
Fluid Mach number
•
Momentum of the released fluid
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•
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considered when engineers design flare pipes (i.e., tail pipes, sub-headers, and header pipes).
Generally, the engineers consider the most governing PSV relieving scenarios when they
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design the flare system. For example, all feasible abnormal scenarios, such as total power failure, external fires, and equipment failure, are considered for designing the size of the flare pipes, size of the PSV orifices, number of PSVs, etc. Among them, the most governing case is considered
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for designing these parameters of the flare system. Furthermore, an additional margin is added to the developed design parameters to cover unanticipated emergency situations. This leads to increased capital costs of the equipment, limited space for equipment installation, and additional weight of the offshore plant (Goyal and Al-Ansari, 2009). Over the past few decades, many researchers have investigated the design of the flare system. For example, Muktikana Sahoo (Sahoo, 2013) developed a steady-state model of the flare system
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in an external fire scenario using industry data and estimated the PSV back pressure and velocity of the relieving fluid. The results indicated that a high flare load in the case of a fire scenario leads to a high back pressure of the PSV, causing severe damage to the PSV. Thus, the study indicated that the diameter of the tail pipe must be increased or the tail pipe should be divided into two separate pipes to reduce the flare load for each pipe. Unfortunately, this study did not
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consider limit of the Mach number that affects the vibration of the pipeline. In another study, Sisshartha Mukherjee (Mukherjee, 2008) numerically optimized flare
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networks in the case of cooling-water failure and external fire scenarios by solving governing
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equations. As one of the constraints for optimization, the maximum values of the PSV back pressure and Mach number in the header were considered. For estimation of these parameters,
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the American Petroleum Institute (API) Standard 521 was employed. During optimization, design variables, including the number of PSVs, configuration of the pipelines, and diameters of
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the pipes, were adjusted, satisfying the proposed constraints. However, only a steady-state model was developed; dynamic behavior, such as changes in the relieving loads over time, was not
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considered. Thus, the design margin was too large for designing pipes with the optimum size for flare networks.
Recently, Prashanth Siddhamshetty (Siddhamshetty et al., 2019) performed a simplified model
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of ignition phenomenon at a wellhead blowout situation. In this study, they reviewed previous papers comprehensively to analyze the relationship between a set of parameters and burning efficiency. Moreover, they developed a simplified model to calculate the burning efficiency, and the results showed high accuracy. After PSV opening, the liquid is generally formed in the pipe due to the depressurization. This liquid formation affects the flame length at the flare stack and
7
flame temperature. In terms of the safety, prediction of flame length and the flame temperature is considered to be quite important. Most flare systems of existing chemical processes were designed according to the steady-state condition. This method uses the sum of flare loads, which are collected in header pipes from different sources of PSVs, to design the diameters of the headers, as described in Figure 2(a).
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However, the flare load that is released from each PSV has a peak point with regard to the flowrate immediately after its release, and the flowrate decreases, as shown in Figure 2(b). Thus,
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if the diameters of the headers are designed according to the estimated maximum flowrates, the
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design margin becomes too large. Additionally, all the main equipment is installed with long distances for safety, and it takes different times for the released fluid from each PSV that is
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installed in the equipment to reach the main header. Consequently, the actual cumulative flare load that passes through the main header is smaller than algebraic sum of the flare loads from the
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PSVs. Therefore, it is necessary to estimate the actual flare loads that pass through the main headers over time from the point of their release until they reach the flare tips. However, few
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studies involving the analysis of flare systems with dynamic simulation have been published.
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(b) Figure 2. Calculation of the total flare load using the (a) steady-state model and (b) dynamic
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model
Gilardi and Tonello (Gilardi and Tonello, 2011) estimated the total flare load along the main
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flare header over time in a power-failure scenario by using a dynamic simulation model. The results indicated that the maximum accumulated flare load that was collected from various
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sources of PSVs was significantly smaller than the algebraic sum of the flare loads from the individual PSV sources. Alban Sirven et al. (Sirven et al., 2011) estimated the total flare load in the case of reflux failure of a crude distillation unit using a DynaSim simulation. The size and number of PSVs were optimized to minimize the total capital cost of the flare system. However, the study focused on the total flare load but did not consider the Mach number of the fluid in the
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(a)
pipes. Ali Shafaghat (Shafaghat, 2016) performed a dynamic simulation of the flare system in an external-fire scenario and analyzed the changes in the fluid properties over time with regard to the pressure, temperature, and mass flowrate during the depressurization period. Then, the estimated Mach numbers from steady-state and dynamic simulation models were compared. The
9
estimated Mach number from the steady-state model exceeded the design criteria, whereas the Mach number from the dynamic model was significantly smaller than the design criteria during the whole relieving period, except for the initial 1 min. Thus, the size of the header pipes could be reduced. However, only the main header was considered in the study. Most previous studies on the design of the flare network system involving dynamic simulation
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have mainly focused on comparing the flare loads, Mach numbers, back pressures, and sizes of the main header pipes in particular emergency scenarios. They seldom revealed the optimum size
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of the whole flare system, including the tail pipes that are connected to PSVs. Furthermore,
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onshore chemical plants were mainly considered by researchers, while its impact on capital cost
2.
Modeling methods
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2.1. General PSV relieving scenarios
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saving is far larger for offshore plants.
In general, a PSV opens owing to overpressure of the equipment, which has various causes.
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Therefore, it is necessary to consider each scenario that causes overpressure of the equipment and apply the results to the design of the flare network system to ensure its safe operation. Relieving scenarios can be divided into two main categories. The first is the continuous-emission
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scenario, which includes purging, control valve leakage, and process venting, and the second is the emergency scenario, in which the flare load is released in a short time for depressurization, e.g., external fires, power failure, and outlet blockages (Davoudi et al., 2014). In general, the emergency scenario is employed to determine the size of the flare network pipes, because the flare relieving load for the continuous-emission scenario is significantly smaller than that for the emergency scenario (Sangsaraki and Anajafi, 2015). The details and design guidelines of the
10
typical emergency scenarios, such as closed outlets, cooling-water failure in the condenser, overfilling, failure of automatic controls, fire, and power failure, are provided in API Standard 521. 2.2. Steady-state modeling A typical flare network system can be modeled in the steady state by using the AFSA. The
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AFSA, which is specialized for the design of the flare network system, can be used to estimate the Mach number and momentum of the fluid in pipes and the noise of the pipes. Furthermore, a
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design engineer can easily verify that the design criteria are satisfied as he/she adjusts the inputs
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of the flare network system. In the AFSA, an engineer must specify inputs, such as the expected relieving load of each PSV, set pressures of the PSVs, orifice sizes of the PSVs, and discharge
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pressures of the PSVs. Once all inputs are entered in the simulation model, the design
estimated. 2.3. Dynamic modeling
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parameters, such as the Mach number, back pressure, and noise level of the pipes, can be
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In contrast to the traditional steady-state model of the flare network system, a dynamic model can analyze the changes in the flare loads and their impact on the back pressure of the PSV and Mach number of the fluid in the pipes over time. Therefore, the estimated results are regarded as
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more accurate and reliable than those from the steady-state model for designing the flare network system. Additionally, it is possible to accurately estimate the uneven distribution of flare loads between PSVs that are installed on the same equipment and to observe the changes in the pressure, temperature, and flowrates under emergency conditions. One of the most important variables in designing the flare network system is the pressure change of the equipment, and the opening tendency of PSVs can be estimated over time according to the pressure change inside a
11
vessel. The factors that affect the pressure of the vessel are the initial operating conditions of the feed stream that enters the vessel (i.e., flowrate, temperature, pressure, and composition), size of the vessel, and closing or opening speed of the valves that are installed in the inlet and outlet pipes of the vessel. In this study, gPROMS Process Builder was used in order to simulate the dynamic state. The gPROMS is based on the dynamic state, and calculate using Equation
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Oriented (EO) method. Therefore, both the flowsheet based model and the custom model are simulated simultaneously. As a result, the calculation speed is much faster than Sequential
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Modulate (SM) method.
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2.4. Estimation of Mach number
The Mach number is defined as the ratio of the velocity of a fluid to the velocity of sound in
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the fluid. In the steady-state model (i.e., AFSA), the Mach number is calculated using Equation
Equation 2.
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1, and the sonic velocity is calculated according to the pressure and density differences using
𝐹𝑙𝑢𝑖𝑑 𝑣𝑒𝑙𝑜𝑐𝑖𝑡𝑦 (1) 𝑆𝑜𝑛𝑖𝑐 𝑣𝑒𝑙𝑜𝑐𝑖𝑡𝑦 𝑖𝑛 𝑡ℎ𝑒 𝑓𝑙𝑢𝑖𝑑 ∆𝑃 (2) 𝑐2 = ∆𝜌 Here, Ma represents the Mach number, c represents the sonic velocity in the fluid, and ΔP and
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𝑀𝑎 =
Δρ represent the differences in the pressure and density, respectively, between the pipe inlet and
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outlet.
In the dynamic model (i.e., gPROMS ProcessBuilder), the Mach number was calculated using
Equation 3, which is described in the API Standard 521. Equation 1 can also be used to calculate the Mach number in the dynamic model; however, API Standard 521 is more widely used by engineers in industry.
12
𝑚̇ 𝑍∙𝑇 (3) 𝑀𝑎 = 3.23 × 10−5 ( )( ) 2 𝑃∙𝐷 𝑀𝑤 Here, 𝑚̇ represents the mass flowrate in the pipe, P represents the pressure of the pipe outlet, D represents the diameter of the pipe, Z is the compressibility factor, T represents the temperature of the pipe outlet, and Mw represents the average molecular weight.
Scenario analysis
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3.
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3.1. Separator outlet blocked discharge
Figure 3. Schematic of the "HP separator outlet blocked discharge” scenario In this study, “high-pressure separator (HP separator) outlet blocked discharge” was selected
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as a governing scenario. The feed stream to an HP separator is directly connected to a riser from the subsea region, and operating pressure of the fluid is very high. Therefore, if the outlet valve of the separator is blocked, the risk of explosion becomes high owing to the high-pressure stream inflow to the equipment. Figure 3 presents a schematic of the “HP separator outlet blocked discharge” scenario. The hydrocarbon mixture passes through the HP separator, and its pressure
13
and temperature are approximately 14.8 bar and 70 ºC, respectively. If the outlet valves are blocked at a certain point, the fluid that continuously enters the vessel accumulates inside the separator, increasing the pressure inside the vessel. Then, as the separator pressure reaches the pressure alarm high high (PAHH) point, the emergency shutdown valve (ESDV), which is connected to the inlet stream, is scheduled to close for approximately 20 s to block the inlet
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stream. While the ESDV closes, the inlet flow continuously enters the HP separator. Finally, the HP separator pressure reaches the set pressure of the PSVs, and the PSVs opens to release the
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fluid inside the vessel and reduce its pressure. The released fluid initially passes through the tail
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pipes of the PSVs, and it gathers in the main flare header pipes. The fluid transported through the
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main header pipes is discharged into the atmosphere after incineration at the flare stack.
Table 1 presents the maximum allowable working pressure (MAWP) and PAHH of the HP
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separator, and Table 2 presents the set pressure of the PSVs installed on the top of the vessel. The sizes of the PSV orifices are identical, whereas the set pressure differs among the PSVs.
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According to API 521, the maximum allowable set pressure of the 1st PSV should be equal to the MAWP, and those of the supplemental PSVs should be <105% of the MAWP. Therefore, the
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highest set pressure was set as 104.5% of the MAWP in this case.
Table 1. Design of the HP separator Item
Pressure [bar]
MAWP
26.634
PAHH
21.690
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Table 2. Set pressure of each PSV Set pressure [bar]
PSV-A
26.634
PSV-B
27.034
PSV-C
27.434
PSV-D
27.834
PSV-E (spare)
26.634
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Item
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3.2. Development of steady-state model
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Figure 4 shows a snapshot of the steady-state model developed using the AFSA for the “HP separator with outlet blocked discharge” scenario. The configuration of the model in Figure 4
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includes PSVs, tail pipes, main header pipes, a flare knockout drum, and a flare stack. The input variables at each PSV, such as the set pressure, orifice size, and mass flowrate of flare load, were
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directly specified in the software. For the modeling of pipes, the design parameters, such as the
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diameter, thickness, materials, and length of each pipe segment, must be specified by a user.
Figure 4. Configuration of the "HP separator outlet blocked discharge" scenario in the AFSA In this study, modulating pilot-operated PSVs identical to those in the actual industrial project were selected. The pilot-operated PSVs mainly consist of a pilot piston and a main valve. The pilot piston presses the main valve through the piston using the vessel pressure. The top area of
15
the pilot piston is larger than the bottom area that is in contact with the main valve. Therefore, the net force presses the pilot piston tightly until the vessel pressure reaches the set pressure. When the vessel pressure reaches the set pressure, the pressure of the pilot piston is vented for reduction of the seating force, and the main valve piston is lifted for depressurization of the vessel. Because the pilot-operated PSV is operated using the vessel pressure, the lift of the main
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valve is not significantly affected by the back pressure (American Petroleum Industry, 2008; Scott and MacKinnin, 2011). For estimation of the relieving load from each PSV, a mass
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flowrate of 267,200 kg/h was specified for each PSV model. For this, the total feed flowrate to
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the HP separator, as shown in table 3, was divided by the number of PSVs, in accordance with the API Standard 521. Two different sizes of tail pipes (10 inches for the 1st tail pipe and 12
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inches for the 2nd tail pipe) were specified in the tail pipe model. A diameter of 24 inches was used for the header pipes, and a diameter of 30 inches was used for the main header pipe
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immediately before the knockout drum. The composition of the relieving fluid (mole fractions) was obtained from the industry project data, as shown in Table 4. The feed condition such as
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flowrate, temperature and pressure of each PSV was assumed by a process engineer, and they were specified as an input for each PSV in the flowsheet. Therefore, the specified feed condition
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of each PSV was referred as relieving condition in a steady-state model.
Table 3. Steady state conditions Relieving conditions Pressure (bar)
26.634
Temperature (ºC)
70
Mass flowrate (kg/h)
1,336,000
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Table 4. Mole fractions under the relieving conditions Mole fraction
Component
Mole fraction
Component
Mole fraction
CO2
0.0029
C6*
0.0041
C15*
0.0032
Nitrogen
0.0019
C7*
0.0064
C16*
0.0026
Methane
0.3490
C8*
0.0062
C17*
0.0023
Ethane
0.0267
C9*
0.0049
C18*
Propane
0.0138
C10*
0.0044
C19*
i-Butane
0.0032
C11*
0.0037
n-Butane
0.0064
C12*
0.0034
i-Pentane
0.0028
C13*
0.0036
n-Pentane
0.0030
C14*
0.0030
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Component
0.0023 0.0022 0.0225
H2S
0.0001
H2O
0.5153
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C20+*
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3.3. Development of dynamic model
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Figure 5. HP separator blocked discharge scenario model in gPROMS
As a next step, the dynamic model was developed using gPROMS ProcessBuilder for the “HP
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separator with outlet blocked discharge” scenario. In contrast to the steady-state model that was developed previously, the main process was included in the model together with the flare network system, as shown in Figure 5. In particular, the dynamic model included a riser stream from the reservoir to the HP separator, the size of the main equipment, and the pressure drops
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through the valves and pipes. Tables 5 and 6 present the operating conditions of the feed stream and its composition (mole fractions), respectively. In contrast to steady-state model, feed condition to PSV is not fixed in dynamic simulation model. In this case, the topside process is integrated with subsea piping system from the reservoir. For the reason, reliving condition of the PSV is estimated through simulation based on the operating condition when an emergency situation occurs. For dynamic simulation of the scenario, an “outlet valve” block was
18
implemented in the model, and it was manually closed once the normal operation was stabilized. The system was stabilized 50 s after the dynamic simulation started. When the pressure of the separator reached the PAHH point (21.690 bar) from its initial point (14.8 bar) owing to the blockage of the outlet valve, the ESDV closed for 20 s. Then, four PSVs, which were installed on the HP separator (excluding the spare PSV), started to open in a row once the pressure of the
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HP separator reached the set pressure of each PSV, and the compressed fluid was released into the tail pipes. As expected, the relieving load changed significantly over time after the PSV
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opened, until it was closed again. Furthermore, the rupture of one PSV affected the other PSVs
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that were connected via the pipes, owing to the interaction of the back pressure and packing
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effect.
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Table 5. Feed stream conditions of dynamic simulation Operating conditions
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Temperature (ºC)
70
Mass flowrate (kg/h)
900,000
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Pressure (bar)
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Table 6. Mole fractions of the feed stream Mole fraction
Component
Mole fraction
Component
Mole fraction
CO2
0.0043
C6*
0.0035
C15*
0.0022
Nitrogen
0.0029
C7*
0.0055
C16*
0.0020
Methane
0.3190
C8*
0.0052
C17*
0.0019
Ethane
0.0273
C9*
0.0031
C18*
0.0013
Propane
0.0137
C10*
0.0036
C19*
i-Butane
0.0027
C11*
0.0028
C20+*
n-Butane
0.0059
C12*
0.0035
i-Pentane
0.0027
C13*
0.0030
n-Pentane
0.0031
C14*
0.0027
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0.0020 0.0022 0.0002
H2O
0.5735
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H2S
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4.
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Component
Results and discussion
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The simulation results of the dynamic model were analyzed. First, the dynamic behavior of the HP separator was analyzed. Then, the dynamic behavior of the flare system, such as the opening
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of the PSVs and the relieving loads from the PSVs, was analyzed. Lastly, the results of the dynamic simulation were compared with those of the steady-state model, and the optimum sizes of the PSV orifices and pipes were estimated according to the sizing criteria of API Standard 521. In this study, the following CPU and GPU were used for computation. CPU: Intel® Xeon® Gold 5120 CPU 2.20GHz
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GPU: NVIDIA GeForce GTX 1080 Ti
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In general, 3 – 5 seconds were taken to perform simulation of the steady-state model, while
200 – 210 seconds for dynamic model. Because the dynamic model considered both the main process and flare network simultaneously, it requires more time for computation. However, the
20
simulation results of the dynamic model showed much higher accuracy compared to the ones from the steady-state model. 4.1. Dynamic behavior of HP separator As shown in Figure 6(a), the HP separator normally operated at a constant pressure of 14.8 bar. Once the gas and liquid outlet valves of the HP separator were blocked at 50 s, as shown in
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Figure 6(b), the pressure of the vessel increased steadily to 27.6 bar (at 80 s) owing to the continuous entrance of the feed stream to the HP separator. For reference, it took approximately
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10 s for the outlet valves to close fully. When the pressure of the HP separator reached the
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PAHH point (21.690 bar) at approximately 67 s, the ESDV closed for approximately 20 s, and the inlet mass flowrate started to decrease rapidly. Because of the continuous inflow of the feed
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stream during the closing period of the ESDV, the pressure continuously increased and reached the PSV-A set pressure (26.634 bar) at 76 s. PSV-A opened rapidly, and a mixture of gas and
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liquids was released to protect the HP separator from further build-up of pressure. Although PSV-A opened, the pressure was accumulated because the opening of only one PSV was
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insufficient to prevent further pressure build-up. Therefore, the pressure slowly increased further and ultimately reached the set pressures of PSV-B (27.034 bar) and PSV-C (27.434 bar). Thus, PSV-B and PSV-C sequentially opened at 77 and 78 s, respectively. On the other hand, PSV-D
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did not open, because the vessel pressure started to decrease afterwards. After the pressure decreased to 26.60 bar, it remained constant below the MAWP (26.634 bar) from 105 s because the SDV was fully closed at 87 s. According to API Standard 521, the maximum allowable overpressure in the case of multiple valve installation should be <116% of the MAWP. Because the highest pressure of the vessel was 27.6 bar, the maximum overpressure was 103.6%, which satisfied the standard.
21
(b)
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(a)
Figure 6. Dynamic behavior of the HP separator: (a) mass flowrate and pressure; (b) opening of
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block valve and ESDV
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4.2. Dynamic behavior of PSVs
Figure 7(a) illustrates the dynamic behavior of the PSVs installed on the HP separator. The
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simulation results indicated that three of the four PSVs (with the exception of the spare) opened
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to protect the vessel. The specified set pressures of PSV-A, PSV-B, PSV-C, and PSV-D were 26.634, 27.034, 27.434, and 27.834 bar, respectively, and the valves opened sequentially over
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time. PSV-A opened at 76 s, PSV-B opened at 77 s, and PSV-C opened at 78 s, and the opened PSVs closed in the reverse order (i.e., C, B, and A). PSV-A released the largest flare load (27.96 kg/s), and PSV-B and PSV-C released smaller flare loads (16.59 and 5.54 kg/s, respectively).
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As described previously, the lift of the pilot-operated PSV was not affected by the back pressure. However, the relieving load of the PSV and the Mach number of the fluid in the tail or header pipes were affected by the back pressure. The back pressure of the PSVs was calculated using the operating pressure of the flare knockout drum, which normally operated at 1 atm, by adding pressure drops through the pipes (Zadakbar et al., 2015; Hernández-Suárez et al., 2007). Figure 7(b) shows the back pressures of PSV-A, B, and C. As PSV-A opened, the back pressures
22
of PSV-B and PSV-C were affected simultaneously. The differences in the flare loads of the
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PSVs and the back pressures caused the differences in the Mach number of the fluid in the pipes.
(b)
(a)
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Figure 7 PSV conditions: (a) mass flowrate; (b) back pressure
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4.3. Mach-number calculation results and comparison with AFSA
A profile of the estimated Mach numbers and the released flare load and outlet pressure of the
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PSV at one of the tail pipes are shown in Figures 8(a) and 8(b), respectively. Once the PSV opened, the Mach number increased rapidly and decreased after reaching the peak point. To
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obtain a conservative design for the sizing of the flare pipes, the maximum values of the Mach
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numbers were considered in this study.
23
(b)
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(a)
Figure 8. (a) Estimated Mach number and (b) released flare load from the PSV and its outlet
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pressure
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In general, the estimated Mach numbers differed between the inlet and outlet points of the same pipe. Because the mass flowrate along the pipe was constant and the pressure was lower at
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the outlet point owing to the pressure drop, the estimated Mach number at the inlet was larger
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the pipe was mainly considered.
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than that at the outlet, as indicated by Equation 3. In this study, the Mach number at the outlet of
The estimated Mach numbers from the steady-state and dynamic models are compared in Figure 9 and Table 7. Overall, the Mach numbers of the dynamic model were smaller than those of the steady-state model, except for pipes 1–4, which were connected to PSV-A and PSV-B.
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The estimated Mach numbers of pipes 1–4 in the dynamic model were larger than those in the steady-state model. Moreover, pipes 1 and 2, which were connected to PSV-A, exceeded the line sizing criteria of 0.7, which can cause safety problems, such as the vibration of the pipe during operation. This is because the PSVs had the same orifice size but different set pressures were specified, which resulted in the release of different amounts of flare loads from the PSVs. Because the largest flare load was released from PSV-A, which had the lowest set pressure, the
24
largest Mach number was estimated in the tail pipes connected to PSV-A. Additionally, the Mach numbers of PSV-B and C, which opened later owing to the high set pressure, were estimated to be lower than that of PSV-A. The Mach number in the tail pipes from PSV-D was
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estimated to be zero because PSV-D was not open, owing to the relatively high set pressure.
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Figure 9. Comparison of the Mach number between the steady-state and dynamic models Table 7. Calculated Mach number of each pipe Tail pipe
PSV-B
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PSV-A
Header
PSV-C
PSV-D
pipe
pipe
pipe
pipe
pipe
pipe
pipe
pipe
pipe
pipe
pipe
pipe
pipe
1
2
3
4
5
6
7
8
12
13
14
15
16
Steadystate
0.6
0.4
0.57
0.37
0.56
0.37
0.57
0.37
0.09
0.19
0.29
0.44
0.27
Dynamic
0.87
0.80
0.74
0.59
0.30
0.21
0.00
0.00
0.00
0.05
0.19
0.47
0.32
25
4.4. Pipe-size optimization results The optimum sizes of the tail and main header pipes were estimated according to the dynamic simulation results and the line sizing criteria of API Standard 521. In practice, the sizes of tail pipes and sub-header pipes should be adjusted to satisfy the maximum value of the Mach number (0.7). For the sizing of the main header, the estimated Mach number should be <0.5. Detailed
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guidelines for the line sizing are presented below (Kerr et al., 2018). Line sizing criteria in API 521: Mach number of header < 0.5
•
Mach number of sub-header and tail pipe < 0.7
•
MABP: 50% of relief-valve set pressure for modulating pilot-operated PSV
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•
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In this study, Equation 3 was used to determine the appropriate size of the pipes by adopting
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the maximum Mach numbers (0.5 or 0.7). The results are shown in Figure 10 and Table 8.
Figure 10. Calculated pipe diameter vs. original design
26
Table 8. Calculated diameter of each pipe Tail pipe Header PSV-A
PSV-B
PSV-C
PSV-D
pipe 2
pipe 3
pipe 4
pipe 5
pipe 6
pipe 7
pipe 8
pipe 12
pipe 13
pipe 14
pipe 15
pipe 16
Original design
10
12
10
12
10
12
10
12
24
24
24
24
30
Dynamic
14
14
10
10
5
6
0
0
0
8
14
24
24
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pipe 1
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As shown in Figure 10, there were many chances to reduce size of the pipes compared with the
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original design. The diameter of most pipes could be reduced or maintained. In the case of tail pipes 1 and 2, which were connected to PSV-A, the diameter must be increased by one or two
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sizes to ensure safe operation. As described previously, the dynamic simulation results proved that larger flare loads were expected to be released through PSV-A than through the other PSVs,
Conclusion
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5.
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whereas in the steady-state model, the same flare load was allocated to each PSV.
A governing scenario of “HP separator outlet blocked discharge” in an offshore plant was adopted to compare the design results of a flare network system between steady-state and dynamic models. For the steady-state model, the widely used software of the AFSA was used,
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whereas gPROMS ProcessBuilder was used to develop the dynamic simulation model. Compared with the steady-state model, the dynamic model estimated the behavior of the relieved fluids over time more accurately, including the flowrates, pressures, and temperatures of the fluids in the pipes. Detailed information regarding the main process, process control scheme, and ratings of major equipment were incorporated into flare network system. Then, the Mach
27
numbers of the fluids in the pipes were estimated for both the dynamic and steady-state simulations, and the results were compared. The results indicated that the original design of the flare network must be modified significantly for capital cost saving and safe operation. Lastly, the optimum sizes of the tail pipes, sub-headers, and main headers were calculated according to both the dynamic simulation results and the standard line sizing criteria from API Standard 521.
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The results indicated that the diameters of most pipes could be reduced, reducing the capital cost. However, the diameters of a few pipes must be increased to avoid vibration and noise of the
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pipes during abnormal operation.
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For sizing of the main header and sub-header pipes, the most extreme case of the emergency situations should be considered as a governing scenario. Because the “HP separator outlet
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blocked discharge” was considered to be the governing case among all potential emergency situations (e.g., power failure, control failure, fire, etc.) in this study, the size of the main header
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and sub-headers should be big enough to to cover all other emergency situations. Furthermore, performance of the control system was examined through dynamic simulation to avoid any
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potential failure of the system.
The refinery, petrochemical and off-shore plants have flare network system to protect the units and humans in abnormal operations. Generally, the flare network system in these plants is
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designed based on the steady-state model. Therefore, high margin is added to the design of flare network system to cover inaccuracy that is caused by the steady-state model. However, as computation speed increases rapidly these days, computational time for dynamic simulation becomes much shortened. For the reason, change of operating condition can be examined more accurately during emergency situation by using dynamic simulation, and we can design more robust and safer flare network system with lower capital cost. This proposed methodology allows
28
us to design the system efficiently with high accuracy. However, it requires substantial information regarding the process and the ratings of the equipment, as well as considerable manhours. Therefore, a new methodology should be developed for designing the flare network system more efficiency while maintaining high accuracy with reduced man-hours.
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Funding This research was supported by a Korea Institute for Advancement of Technology (KIAT) grant
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funded by the Korean Government (MOTIE) (P0008475, Development Program for Smart
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Digital Engineering Specialist). It was also supported by the Basic Science Research Program through the National Research Foundation of Korea (NRF) funded by the Ministry of Education
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(2019R1F1A1058979).
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Declaration of interests
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The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.
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Bader, A., Baukal, C. E., Bussman, W., 2011. Selecting the proper flare systems. Chem. Eng. Prog., 107(7), 45–50.
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Bjeere, M., Eriksen, J. G. I. ., Andreaden, A., Stegelmann, C., Mando, M., 2016. Analysis of Pressure Safety Valves for fire protection on offshore oil and gas installations. Process. Saf.
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Environ. Prot., 105(Iso 23251), 60–68. doi:10.1016/j.psep.2016.10.008 Chopinet, F., Marcec-Rahelic, N., 2015. Upgrade of a Flare Network for a New Refinery Unit.
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Goriva I Maziva, 54(1), 32–43.
Davoudi, M., Aleghafouri, A., Safadoost, A., 2014. Flaring networks assessment in South Pars
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Gas processing plant. J. Nat. Gas. Sci. Eng., 21, 221–229. doi:10.1016/j.jngse.2014.08.008 Gilardi, C., Tonello, M., 2011. Dynamic simulation, HIPS & combined probability analysis reduce flare loads. Associazione Termotecnica Italiana(ATI), Milano, Italy, March 31, 2011.
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Goyal, R. K., Al-Ansari, E. G., 2009. Impact of emergency shutdown devices on relief system sizing and design. J. Loss. Prev. Process. Ind., 22(1), 35–44. doi:10.1016/j.jlp.2008.07.008
Hernández-Suárez, R., Puebla, H., Aguilar-López, R., 2007. Parametric approach for the optimal design of knockout drums. Ind. Eng. Chem. Res., 46(21), 7008–7017. doi:10.1021/ie0701930
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