FIRE SURVIVAL OF PROCESS VESSELS CONTAINING GAS J. Nylund * The present work is a theoretical evaluation of the ability of process vessels to survive hydrocarbon fires when the vessels are designed and equipped with pressure relief valves and depressuring valves according to current codes and standards. Acceptable delay times for depressurization activation after a hydrocarbon fire has ensued are established for normal and controlled depressurization techniques. A relationship between impinging hydrocarbon jet size and the time period to plastic instability and subsequent rupture is established. INTRODUCTION Local jet and pool fires can develop into major catastrophies if not controlled at an early stage. Loss of structural integrity and containment can accelerate the development of such accidents. Rupture of process vessels can develop into fire-balls with high damage potential. To control these risks process vessels are designed according to codes and standards. These codes and standards have their basis in certain philosophies which are expressed through design equations. To be useful these design equations have to be rather simple. In the present study the adequacy of such simple equations is investigated through more rigorous analyses. A process vessel is designed and equipped according to recognized codes and subjected to fire loads. Failure criteria are proposed, and the ability of the process vessel to survive the chosen fires scenarios is investigated. * Det norske VERITAS, N1322 - Hoevik, Norway 137
PROCESS VESSEL/VALVE DESIGN Process Vessel The process vessel in this study has been designed according to BS5500, /1/. The proposed failure criteria require data for the effect of increased material temperature (T m) on Youngs modulus (E) and the yield strength (o y). Pressure Relief Valves The pressure relief valve is designed for fire relief according to API /2/. The fire load is 50% and 100% of the vessel exposed to fire. Depressuring Valve The valve for normal depressuring is designed according to a simplified method by Grote /3/. According to this method the throat area of the depressuring valve can be calculated from the equation Paruit and Kimmel /4/ have proposed a method for controlled depressuring. The benefit of this method is according to the authors to obtain a constant mass flow throughout the 15 minutes depressuring and in this way reduce the maximum relief amounts to the flare stack. For the subject vessel the constant mass rate is found to be 4.724 kg/s. Table 1 shows the data for the process vessel, the pressure relief valves and the depressuring valve. TABLE 1 Process vessel diameter 3.0 m Total length of vessel 13.0 m Vessel shell thickness 53.0 mm Yield strength of vessel material (API-5LX-60) 420 N/mm 2 Design stress at 90 centigrades 245 N/mm 2 Design pressure (max. allowable working pressure) 8.28 MPa Working pressure 8.2 MPa Working temperature 300 K Relief valve fire relief set pressure 10.0 MPa Discharge coefficient 0.975 Throat area at 50% fire exposure 2.016. 10 4 m 2 Throat area at 100% fire exposure 4.031. 10 4 m 2 Depressuring valve discharge coefficient 0.74 Throat area 1.153. 10-3 m 2 138
PROCESS ANALYSIS TOOLS To analyse the process vessel with pressure relief and depressuring valves when exposed to fire loads, the computer program BLOW-DOWN /5/ has been used. This is a modular computer program for analysis of hydrocarbon gas flow. The progam is capable of handling real hydrocarbon mixtures, but in this study pure methane gas has been applied. Process Vessel The element program VESSEL solves the 1st Law of Thermodynamics for gas in the vessel and the vessel shell. External convection and radiation heat transfer are accounted for as well as internal convection and radiation. Heat transfer coefficients are continuously updated throughout the calculations. Valves The element program VALVE analyse the flow through the pressure relief and the depressuring valves. The element is adiabatic and is capable of handling both sonic and sub-sonic flow. Vessel Failure Criteria The proposed failure criterion for local jet fires impinging on a pressure vessel is an application of the work performed at Battelle's Columbus Laboratories regarding defects in pressurized pipes and vessels. According to Kiefner et al /6/ the stress concentration factor due to a longitudinal through-wall flaw in a vessel can be expressed as Stress concentration factor due to surface flaws can be expressed in terms of M T and the flaw depth - shell thickness ratio (δ/t): Rupture of the vessel takes place when the hoop stress concentration (σ θ M s) reaches the flow stress which accounts for strain-hardening and can be expressed as yield strength plus a constant. This criterion applies to ductile materials and also surface flaws due to corrosion. In such cases parabolic shape of the defect (s = 2/3) can be used up to ( ) = 4.5 over which rectangular shape (s = 1) has to be used. In case of a local jet fire, the heat affected material loses its ability to carry loads, and a situation similar to a surface flaw exists. The equivalent (δ/t) ratio, as seen from the unheated surrounding material, can be expressed as where E o denotes the Youngs Modulus for the unheated material. 139
By use of Equations (4) through (6) and the effective jet fire diameter as surface flaw length, the failure pressure as function of time can be calculated. Strain-hardening is disregarded to account for some heat conduction along the vessel shell. Above a certain size a surface flaw has to be considered global and the flaw depth δ as general thickness reduction. Brakestad and Wiik /7/ suggest that surface flaws where ( )> 8 are to be taken as global surface flaws. Global failure criterion occur when the vessel shell is fully plastified. In case of global fires the failure criterion can be expressed as: The pressure above which the vessel is rupturing, the failure pressure, is calculated in element program TEMP, which also calculates the temperatures through the vessel shell. FIRE SCENARIOS The fire scenarios considered are jet fires of various size impinging on the side of the process vessel and a pool fire totally engulfing the process vessel. Both scenarios have relevance to offshore and onshore industry. The effective jet fire diameter D J is used in this study to account for several types of jets. The diameter of the undisturbed jet will heavily depend on whether the jet is burning before it hits the vessel or not, and whether the jet is underexpanded or not. The flame temperature is taken to be 1500 K for both the jet and the pool fires. The connective heat transfer coefficient in case of impinging jets is taken to be 250 W/m 2 K. Wesson /8/ recommends 28 W/m K as convective heat transfer coefficient for pool fires. Jet Fires FIRE RELIEF Figure 1 through 4 show the results from the analyses of the process vessel subjected to fire jets of varying effective diameter D J. For D J = 1 m (Fig.1) we see that the pressure p v and the temperature T g in the vessel increase just slightly since the fire exposed area is relatively small. The failure pressure p f is above the vessel pressure throughout the 15 minute period and no vessel rupture will thus occur. For effective fire jet diameters 2 and 3 meters, Figs. 2 and 3, we see that the vessel pressure increase more with time, and that the vessel ruptures after about 7.5 and 6.5 minutes respectively. Figure 4 shows that the vessel will rupture after about 4.5 minutes when the vessel is subjected to a jet fire of effective diameter 4 meter. The fire jet in this case is so large that the failure criterion for global fires is used. 140
Common for the results shown for D J = 2, 3 and 4 meters is that the pressure relief valve represents no protection for the vessel since rupture occurs at pressures far below the set pressure. Pool Fires In case the process vessel is fully engulfed in a fire, the vessel pressure increase much quicker than in case of local jet fires. Figure 5 shows the pressure in the process vessel for the cases when the pressure relief valve is not working, designed for 50% fire exposure (p vr50%), and designed for 100% fire exposure (p vr100%). If we for a moment disregard the structural strength of the vessel, we see that the pressure relief valve designed for 100% fire exposure has too little capacity to maintain to maintain the vessel pressure at 10 MPa. However, the vessel pressure exceeds the failure pressure after about 5 minutes, and a slightly higher relief valve capacity would not change this. From the figure we see that the pressure relief valve does not represent any protection of the vessel, and that the vessel would rupture less than one minute earlier if no pressure relief valve were fitted. FIRE DEPRESSURING According to the API code (2) the process vessel is to be depressurized down to 0.79 MPa within 15 minutes. In the literature there has been a tendency of focusing on the depressuring time, while the underlaying philosophy is that the vessel pressure has to be reduced quicker than the loss in mechanical strength due to increased shell temperature. Neither Grote /3/ nor Paruit and Kimmel /4/ take into account the fire situation, and the ability of their approach to cope with "cold" depressuring and depressuring in presence of fire is analysed. Normal Depressuring Figure 6 shows the results from normal depressuring from the working pressure in cold conditions through the depressuring valve designed according to Grote /3/ and shown in Table 1. The vessel pressure decreases nicely to below 0.79 MPa within 15 minutes. The mass flow rate m decreases from initial 12.9 kg/s to 1.2 kg/s after 15 minutes. Figure 7 shows the depressuring from the same working pressure, but in a totally engulfed fire condition. Initially the vessel pressure drops as in cold conditions, but due to the heat input, not as fast. The final vessel pressure after 15 minutes is above 0.79 MPa, but safely below the failure pressure throughout all the depressuring period. Depressuring valves designed according to Grote /3/ is thus capable of safe depressuring of a vessel when the start pressure is the working pressure. In Figure 5 we saw that vessel rupture occurred after about 5 minutes and that the vessel pressure at that instant was 10.2 MPa. We would like to investigate if a 5 minutes delayed depressuring could protect the vessel. Figure 8 shows this situation where p yd denotes the vessel pressure after the delayed depressuring. We see that the vessel pressure during depressuring is kept below the failure pressure p f and can conclude that up to 5 minutes delayed depressuring can be acceptable without vessel failure to occur. 141
Controlled Depressuring The depressuring control system proposed by Paruit and Kimmel /4/ to obtain constant mass rate during depressuring is shown in Figure 9. The depressuring takes place through a restriction orifice RO and a pressure control valve PCV in parallel. The PCV is controlled by a pressure indicator controller PIC sensing the flare header backpressure. At start of the depressuring the PCV is closed and the PCV flow area is increased during depressuring so that the flare header backpressure is kept constant. We have assumed that a controlled depressuring system which maintain constant mass rate is fitted to the subject vessel and cold depressuring from working pressure is performed. Figure 10 shows that constant mass rate 4.724 kg/s results in pressure decrease nicely to below 0.79 MPa within 15 minutes. The same vessel but totally engulfed in fire is shown in Figure 11. We see that a constant mass rate 4.724 kg/s is not sufficient to keep the vessel pressure below the failure pressure. The vessel will thus rupture about 7 minutes after the depressuring has started, and controlled depressuring method maintaining constant mass rate is thus not an acceptable protection method for the subject vessel. It is evident that delayed depressuring will not work either. The depressuring control system in Figure 9 maintains the flare header pressure drop constant, but the gas temperature, and thus the gas density will be sensitive to whether the vessel is exposed to fire or not. To check the ability of the depressuring control system a simplified analysis of the flare header is performed. The pressure in the flare header is rather low, and we assume that the pressure drop is mainly due to friction and that the D'Arcy equation can be applied. D'Arcy friction factor can be expressed as Equations (8) and (9) and the continuity equation give that the mass rate in case of constant flare header backpressure can be expressed as A typical flare header backpressure 0.15 MPa is applied, and controlled depressuring in cold conditions and under total fire engulfment are analysed. With initial mass rate 4.724 kg/s the change in mass rates is shown in Figure 12. We see that the proposed depressuring control system will increase the mass rate during cold depressuring when the flare header backpressure is kept constant. This is mainly due to the gas temperature which decrease from initially minus 9 centigrades (264K) to minus 121 centigrades (152K) downstream the depressuring control system and thus increase the gas density. Relative to Figure 10 the pressure drops to below 0.79 MPa after about 13 minutes. In case of total fire engulfment we see from Figure 12 that the mass rate is reduced from initially 4.724 kg/s to less than half this value at the end of the depressuring. The gas temperature downstream the depressuring control 142
system does in this case increase from initially minus 9 centigrades (264K), as for the cold depressuring, to plus 646 centigrades (919K) which will reduce the gas density. The time to vessel rupture will be just slightly shorter than shown in Figure 10. CONCLUSION The results from the analyses performed in this study can be summarized as follows. A process vessel subjected to fire can rupture as early as 4 to 5 minutes after the start of the fire. Both the jet fire (Figure 4) and the pool fire (Figure 5) give the same result. Common for these two fire scenarios is also that the pressure relief valve provides no real protection of the process vessel. Normal depressuring technique as represented by Grote provides a real protection of the pressure vessel when activated within 4-5 minutes after start of the fire. This applies to both jet fires and pool fires. Controlled depressuring technique based on constant mass flow rate or constant flare header backpressure throughout the depressuring time provides no real protection in fire conditions when fitted to a single gas process vessel. It should be stressed, however, that the results from this study are not universal, and other conditions may call for specific analysis. 143
A D C c D D D J d E f NOMENCLATURE depressuring valve cross section area proportionality constant discharge coefficient vessel diameter effective jet fire diameter flare header diameter Youngs modulus D'Arcy friction factor L flaw length along vessels axis 1 length of flare header M S surface flaw stress concentration M T m P f p v P vr P vd P 1 P 2 p Re s T g T m t u v Z δ σ y σ θ τ ρ through-wall stress concentration mass flow rate vessel failure pressure vessel pressure vessel pressure at relief vessel pressure at depressuring vessel pressure before depressuring vessel pressure after depressuring flare header pressure drop Reynolds number surface flaw shape factor gas temperature vessel shell temperature vessel shell thickness gas velocity vessel volume gas compressibility flaw depth yield strength hoop stress depressuring time gas density µ gas dynamic viscosity 144
REFERENCES 1. British Standard 5500 "Specification for Unfired Fusion Welded Pressure Vessels" 2. API RP 520 I & II "Recommended Practice for the Design and Installation of Pressure-Relieving Systems in Refineries" American Petroleum Institute 3. Grote, S.H.: "Calculating Pressure Release Times" Chemical Engineering July 17, 1967 4. Paruit, B. and Kimmel, W.: "Control Slowdown to the Flare" Hydrocarbon Processing October 1979 5. Nylund, J.: "Element Program System BLOW-DOWN General Description" Det norske VERITAS 6. Kiefner, J.F., Maxey, W.A., Eiber., R.J. and Duffy, A.R.: "Failure Stress Levels of Flaws in Pressurized Cylinders" Progress in Flaw Growth and Facture Toughness Testing, ASTM STP 536 American Society for Testing and Materials 1973, pp 461-481 7. Brakestad, H. and Wiik, T.: "Riser Rupture, Causes and consequences - Part report II: The Influence of surface flaws on the strength of pressurized steel pipes" VERITAS Report No.79-0303 8. Wesson, H.R.: "Considerations Relating to Fire Protection Requirements for LNG Plants" 145
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