API Standard 521 New Alternative Method to Evaluate Fire Relief for Pressure Relief Device Sizing and Depressuring System Design 2014 Journal of Loss

Journal of Loss Prevention in the Process Industries 27 (2014) 21e31

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API Standard 521 new alternative method to evaluate fire relief for pressure relief device sizing and depressuring system design Edward Zamejc EZ Relief Systems Consulting, Inc., 7905 Grady Circle, Castle Rock, CO 80108-6112, USA

a r t i c l e i n f o

a b s t r a c t

Article history: Received 31 October 2012 Received in revised form 28 October 2013 Accepted 29 October 2013

Since the 1950’s, API Standards have provided guidance on determining relief loads for equipment exposed to pool fires. The API method is empirical based on tests performed in the 1940’s. There is increasingly widespread interest in analytical methods based on heat transfer principles to model fire heat input. The API committee agreed to include an analytical method in the 6th edition of API Standard 521 to establish relief loads for pressure relief devices and to design depressuring systems for the fire scenario. The analytical method provides more flexibility than the empirical method but has limitations (e.g., too many permutations are possible leading to potential under-sizing of the pressure relief device). This paper discusses the basis for the empirical method in API Standard 521 and provides comparisons of the empirical and analytical method with two more recent large-scale pool fire tests. This pool fire test data indicates that the empirical method will provide a conservative estimate of pool fire heat input for most applications and is still the method of choice when designing pressure relief systems. However, these recent tests indicate the empirical method needs to be modified when a vessel or equipment is partially confined by adjacent embankments or walls equal or greater than the vessel height. In such cases, the wetted area exponent should be 1.0 instead of 0.82. The analytical method is useful in determining time-versus-temperature profiles for heating unwetted vessels of varying wall thicknesses and materials of construction. These profiles, which depend upon the type of fire (e.g., unconfined pool fire, jet fire, etc.), can be combined with tensile strength and stressrupture data to specify a depressuring system’s pressure-versus-time profile. This will minimize failure and/or mitigate the effects of failure due to overheating from fire exposure. Ó 2013 Elsevier Ltd. All rights reserved.

Keywords: Pool fire exposure Jet fire exposure Vessel failure Tensile strength Time to failure Fire heat input

1. Introduction American Petroleum Institute (API) Standard 521 “Pressure Relieving and Depressuring Systems” is an internationally recognized engineering standard used to design pressure relief systems, disposal systems (e.g., flares), and depressuring systems (ANSI/API Standard 521, 2013). It is continually being reviewed, with new editions published in about 5-year intervals. A technical committee consisting of industry representatives, engineering contractors, and regulators recommend and integrate modifications into the Standard. These modifications involve lessons-learned from incidents or near-misses, advances in engineering methodologies, and new guidance based on shared experiences of the members or inspired by technical inquiries. The new API 521 6th edition includes an analytical method to establish relief loads for pressure relief devices and to design E-mail address: [email protected]. 0950-4230/$ e see front matter Ó 2013 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.jlp.2013.10.016

depressuring systems for the fire scenario. The analytical method will complement, but not replace, the existing empirical method. It is important to establish the scope of API 521 and differentiate it from API 2000. API 521 covers pressure vessels and processing equipment (e.g., vessels design in accordance with ASME Section VIII, Division 1 and similar pressure vessel design codes). In contrast, API 2000 covers low pressure, atmospheric, and refrigerated storage tanks designed in accordance with storage tank standards such as API 650. The current pool fire heat input equations in API 2000 are the same as those in NFPA 30. They were established in a 1963 meeting between API and NFPA and are based on a fire test and experience with storage tank fires. Because the origin/basis of API 521 and API 2000 fire equations are different and the scope of the equipment design codes are different, the fire exposure guidance API 521 and API 2000 can neither be interchanged nor compared (i.e., use API 521 for pressure vessels and API 2000 for storage tanks). The subsequent discussion relates only to API 521.

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2. Fire scenario e API empirical method 2.1. Basis of the API empirical method to evaluate pool fires The fire scenario generates the most technical inquiries of any topic in API 521. The current method, given in Equations (1) and (2), is an empirical method based on fire tests performed in the 1940’s. Pool fire heat input with adequate drainage and prompt firefighting:

Q ¼ C1 $F$Aws 0:82

(1)

Pool fire heat input without adequate drainage and prompt firefighting:

Q ¼ C2 $F$Aws

0:82

(2)

where: Q is the total heat absorption (input) to the wetted surface, expressed in W (Btu/h); C1 is a constant [ ¼ 43,200 in SI units (21,000 in USC units)]; C2 is a constant [ ¼ 70,900 in SI units (34,500 in USC units)]. F is an environment factor for fireproofing (F ¼ 1 for no fireproofing); Aws is the total wetted surface, expressed in square meters (square feet). Note 1: the SI equation constants include a conversion factor for (Aws)0.82. Note 2: Inadequate drainage implies the burning liquid can engulf the vessel resulting in increased heat input as compared with a non-engulfing type pool fire where the burning liquid drains away from the vessel. Note 3: Firefighting reduces the fire heat input to a vessel by water spray cooling of vessel surfaces. The origin of the empirical method can be traced back to the 1950’s, when the API Pressure Relief Systems (PRS) technical committee analyzed the available pool fire test data and developed empirical equations to determine the pool fire heat input to a vessel (Heller, 1983). This heat input could then be used to calculate the fire relief load by dividing by the heat of vaporization. These empirical equations include the:  Maximum fire heat input (i.e., maximum heat flux absorbed by the vessel and its contents)  Effect of wetted surface area of the vessel or equipment (i.e., area of the equipment in contact with liquid or below liquid level) on fire heat input  Effect of burning liquid drainage (i.e., whether the pool fire engulfs the vessel) on fire heat input The maximum fire heat input into a vessel is sometimes confused with the flame surface heat flux (i.e., pool fire heat duty divided by flame surface area) and the incident heat flux at a vessel exposed to the fire. The incident heat flux excludes reduction in heat flux due to the absorptivity of the vessel and re-radiation from the vessel. Based on plotting hydrocarbon pool fire test data (see Fig. 1), the maximum heat input into the vessel was determined by the API committee to be 34,500 BTU/h ft2 of wetted surface (see constant C2 in Equation (2)). This maximum heat input would occur when the wetted surfaces of the vessel are completely and continuously exposed to flame. Outdoor pool fires are easily influenced by even relatively calm wind conditions; wind causes flames to move around, thereby only intermittently exposing surfaces of larger vessels to the highest incident heat flux. To determine the effect of vessel wetted area, the

Fig. 1. Pool fire heat input versus wetted area exposure. Test data sources: a) API project test No. 1, b) API project test No. 2, c) Rubber Reserve Corporation test No. 17, d) Standard Oil Company of California, e) Underwriters Laboratories, Inc. (ANSI/API Standard 521, 2013 Table A.1).

API Committee plotted the total heat input versus the wetted area from several pool fire tests and an actual pool fire (see API 521 Table A.1). The results, shown in Fig. 2, indicate the heat input correlates with a 0.82 exponent on the wetted area (Aws). It should be noted that, per convention, the heat input across only the wetted surfaces of vessels containing a liquid that can boil is considered when designing pressure relief systems for the fire scenario using the empirical method. The effect of heating unwetted surfaces and gas-filled vessels is discussed later in this article. The effect of drainage was determined from Hottel’s pool fire test data (ANSI/API Standard 521, 2013, Table A.1, Test 1 and Test 2). Test 1 actually consists of the average of 31 tests without either drainage or firefighting. Test 2 consists of the average of 8 tests with drainage and 5 tests with both drainage and firefighting. The ratio for the pool fire heat input from Test 2 to Test 1 (no drainage nor firefighting) is 17,400/30,500 ¼ 0.6. Hence, the maximum pool fire heat input to vessels with adequate drainage and firefighting is 34,500*0.6 ¼ 21,000 BTU/h ft2 (see constant C1 in Equation (1)). Hottel’s test data from Personal correspondence (1950), shown in Table 1, suggest that drainage alone (no firefighting) has a comparable reduction in heat input. The remaining parameter of the empirical method, designated “F”, is the environment factor which credits for adequate fireproofing (See ANSI/API Standard 521, 2013). Fireproofing that meets the requirements of API 521 reduces the pool fire heat input to a vessel thereby reducing the relief requirements. Fireproofing will also reduce the vessel wall heatup rate thereby increasing the

Fig. 2. API 521 Table A.1 fire tests e Wetted area versus fire heat input. Data sources: all tests shown in ANSI/API Standard 521 (2013) Table A.1 as well as the actual plant fire involving a 380 butane sphere.

E. Zamejc / Journal of Loss Prevention in the Process Industries 27 (2014) 21e31 Table 1 Hottel pool fire test data showing effect of drainage with and without firefighting (Personal correspondence, 1950). Run

Drainage

Firefighting

Average heat input, BTU/h ft2a

10 11 12 13 14 15 17 19 18

Yes Yes Yes Yes Yes Yes Yes Yes Yes

22,862 30,081 9819 19,517 21,707 26,953 2166 16,436 6594

22

Yes

23

Yes

24

Yes

25

Yes

None None None None None None None None Chemical foam for 30 s then mechanical foam Chemical foam for 30 s then mechanical foam Chemical foam for 30 s then mechanical foam Chemical foam for 30 s then mechanical foam Chemical foam for 30 s then mechanical foam Average for all tests Average for tests with drainage, but no firefighting Average for tests with both drainage and firefighting

a

9337 14,535 25,028 20,672 17,362 18,693 15,233

Heat input into wetted surface area.

23

Pressure relief systems designed for fire exposure require a total heat input into the relevant surfaces (e.g., wetted areas). Because the fire heat intensity will vary with time and location, the total heat input should be determined using a “surface average heat flux” obtained by averaging the fire heat intensity across the entire flame volume. In contrast, when designing depressuring systems, the peak fire heat intensity (designated as the “local peat heat flux”) is important because localized overheating in a small area can result in equipment failure due to overheating. It should be noted that API 521 5th and prior editions provide design guidance for pressure relief and depressuring systems only for open pool fires. Also, the 34,500 Btu/h ft2 maximum fire heat input used in the API empirical method includes the vessel absorptivity, which can reduce the incident heat flux by 20e70%. It is important to note that in the case of jet fire exposure of vessels, a pressure relief device will usually not provide significant protection/mitigation benefit because the vessel walls can be heated by an impinging jet fire to a temperature where the material loses strength resulting in vessel failure. In these cases, a properly designed depressuring system is one potential mitigation alternative that can be used. 3. Analytical method to evaluate fires

. Open pool fire e 16,000e48,000 Btu/h ft2 (50e150 kW/m2) . Confined pool fire e 32,000e79,000 Btu/h ft2 (100e250 kW/m2) . Jet fire e 32,000e127,000 Btu/h ft2 (100e400 kW/m2)

Several authors recommended using an analytical method to model the fire scenario to overcome limitations with the empirical method and to provide more flexibility in modeling (e.g. Energy Institute, 2003; Roberts, Medonos, & Shirvill, 2000; Salater; Salater, Overa, & Kjensjord, 2002; SCANDPOWER, 2004; Shirvill). The analytical method is based on theoretical heat transfer equations that include radiative and convective heat transfer terms. In contrast, the API 521 empirical method determines the fire heat input based on a correlation derived from pool fire test data where all of the heat transfer terms and effects are lumped together as shown in Equations (1) and (2). Because the analytical methods are becoming more widely used, the API 521 committee agreed to incorporate the analytical method as an alternative to the empirical method and to include guidance on its application. The empirical method will still be recommended as the preferred method to evaluate most fire scenarios involving pressure relief system design with the analytical method preferred for special cases and fires outside the scope of the empirical method. The analytical method to evaluate fires is a basic heat transfer equation as shown in Equation (3). The method determines the fire heat input to a vessel and conservatively ignores internal heat transfer limitations. It can be applied to all types of fires including open pool fires, confined pool fires, and jet fires. The proposed typical ranges in the parameters are given in Table 2 for the surface average heat flux for pool fires. Recommended values, which should be used where data or other resources are unavailable, are given in ANSI/API Standard 521, 2013. ANSI/API Standard 521, 2013 also provides typical values of parameters for the local peak heat flux of pool fires and for the local peak and surface average heat fluxes for jet fires.

These peak fire intensities generally correspond to locations within the fire where the stoichiometric fuel-to-air ratio is equal to one. Because of the effects of ventilation (e.g., wind effects and confinement), fuel type, fuel-air stoichiometry and other factors, the peak fire heat intensity is generally only observed in localized parts of the flame volume. Most of the flame volume has significantly lower fire intensities than the peak. Note that API 521 does not provide specific guidance on completely confined fires because of their complexity (e.g., sensitivity to ventilation effects) which generally requires case-by-case evaluations.

where: qabsorbed is the absorbed heat flux from the fire, expressed in Btu/ h ft2 (W/m2); s is the StefaneBoltzmann constant ¼ 0.1713  108 Btu/ h ft2  R4 (5.67  108 W/m2 K4); asurface is the equipment absorptivity, dimensionless; εfire is the fire emissivity, dimensionless;

time to failure of gas-filled vessels. The 6th edition has an expanded expression to account for multi-layer fireproofing. It is important to note that the API 521 criteria for determining adequacy of fireproofing correspond to typical pool fire exposure, but not jet fires because of the erosive power of the momentum jet. API 521 does not allow credit for fire protection systems because of reliability aspects. 2.2. Limitations of the API empirical method to evaluate fires There are two types of fires relevant to pressure relief and depressuring system design e pool fire and jet fire. A pool fire is defined as a “burning pool of liquid.” A jet fire is a “fire created when a leak from a pressurized system ignites and forms a burning jet.” A pool fire can be classified as an open pool fire, a confined pool fire, or somewhere in between. A confined pool fire is defined as a “fire inside a building or a compact process module where the walls and/or surrounding equipment can reradiate and preheat the combustion air causing higher heat fluxes than an unconfined (i.e., open) fire.” Generally only pool fires are considered when designing pressure relief systems while both pool fires and jet fires are often considered when designing depressuring systems. Typical ranges for peak fire heat intensity (i.e., incident heat flux) are (Energy Institute, 2003):


4 4 qabsorbed ¼ s$ asurface $εfire $Tfire  εsurface $Tsurface 
þ h$ Tgas  Tsurface

(3)

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Table 2 Typical range in analytical method (Equation (3)) parameters for an open pool fire “surface average heat flux”. Parameter

Description

Pool fire surface average heat flux parameter range

εfire εsurface

Hydrocarbon flame emissivity Equipment emissivity Equipment absorptivity Convective heat transfer coefficient between equipment and surrounding air Temperature of combustion gases flowing over the surface Fire temperature

0.6e1.0 0.3e0.8 0.3e0.8 1.76e5.28 Btu/h ft2  R (10e30 W/m2 K)

asurface

h

Tgas Tfire Tsurface

Equipment temperature StefaneBoltzmann constant

qfire

Fire heat flux e a wider range is possible Absorbed heat flux at start of the fire

s

qabsorbed

1392e2112  R (932e1652  F) 773e1173 K (500e900  C) 1572e2292  R (1112e1832  F) 873e1273 K (600e1000  C) Increases as surface heats up 0.1713  108 Btu/h ft2  R4 (5.67  108 W/m2$K4) 9510e31,700 Btu/h ft2 (30e100 kW/m2) 7925e23,775 Btu/h ft2 (25e75 kW/m2)

εsurface is the equipment emissivity, dimensionless; Tfire is the fire temperature, expressed in  R (K); Tsurface is the equipment temperature, expressed in  R (K); Tgas is the temperature of air/fire in contact with the equipment surface, expressed in  R (K); h is the convection heat transfer coefficient of air/fire in contact the equipment, Btu/h ft2  R (W/m2 K); s$asurface$εfire$T4fire is the radiative heat flux to the equipment; s$εsurface$T4surface is the re-radiation from the equipment; h$(Tgas  Tsurface) is the convection heat transfer between the combustion gases and the equipment’s surface. h is the convection heat transfer coefficient of air/fire in contact the equipment, Btu/h ft2  R (W/m2 K); s$asurface$εfire$T4fire is the radiative heat flux to the equipment; s$εsurface$T4surface is the re-radiation from the equipment; h$(Tgas  Tsurface) is the convection heat transfer between the combustion gases and the equipment’s surface. Note that published heat transfer coefficients and emissivities are often empirically determined but the test conditions may not accurately represent the conditions associated with a particular fire. Caution must be taken when specifying the parameters because a wide range in fire heat inputs can result. When applying the analytical method to sizing pressure relief devices, the total heat input into the vessel shall use the wetted area to the 1.0 exponent, not the 0.82 exponent used in the API empirical method as shown in Equations (1) and (2). The 0.82 exponent is empirically derived from fire test data as shown in Fig. 2 and has no theoretical basis. It is important to note that API 521 provides methods to determine the fire heat input to equipment while it is up to the user to determine the vessel area exposed to a fire based on the type, size, configuration and location of their postulated fire. Consequence modeling is outside the scope of API 521. Application of the analytical model to pool fires is discussed in Section 4.0 Application of the analytical method to modeling jet fires is given in Salater and Overa, (2004).

gallon (125 m3) LPG tank car was exposed to a semi-enclosed pool fire where the pool fire and tank car were located in a pit, with embankments on all sides exceeding the tank car height (no roof). Fire and wall temperatures versus time at the top of the front and rear walls of the tank car are shown in Fig. 3a and b, respectively. The analytical method (Equation (3)) was used by the author in an attempt to reproduce the wall temperature versus time. The reported values in Anderson et al., 1974 were used for the parameters where given (e.g., fire temperature, fire emissivity, initial temperature). There is insufficient technical basis to theoretically predict the value of the unknown parameters that would apply to this specific fire due to the fire variability. Hence, values for the unspecified parameters were adjusted by trial-and-error until the calculated timeetemperature profile approximated the wall temperature data from the fire test shown in Fig. 3b. The values of parameters selected to model tank car rear wall temperature versus time at several locations were: εfire ¼ 0.62 (determined by BRL) εsurface ¼ 0.5 asurface ¼ 0.5 h ¼ 1.76 Btu/h ft2  R (10 W/m2 K) Tgas ¼ 2112  R (1652  F) 1173 K (900  C) Tfire ¼ 2112  R (1652  F) 1173 K (900  C) e see Fig. 3a Tsurface ¼ Initially @ 529  R (69  F) 294 K (21  C)

4. Application of the analytical method to model recent pool fire tests 4.1. Use of the analytical method to model the Ballistics Research Laboratory (BRL) pool fire test wall temperature versus time In 1974, a pool fire test was performed by the Ballistics Research Laboratory (Anderson, Townsend, Zook, & Cowgill, 1974). A 33,000

Fig. 3. (a). Ballistics Research Laboratory pool fire test data illustrating fire temperature versus time at the top of the front and rear walls of a rail tank car. (b). Ballistics Research Laboratory pool fire test data illustrating rail tank car wall temperature versus time at the top of the front and rear walls.

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qfire ¼ Maximum of 16,700 Btu/h ft2 (52.7 kW/m2) (calculated) qabsorbed ¼ Maximum of 9560 Btu/h ft2 (30.2 kW/m2) (calculated) qfire is calculated by setting εsurface ¼ 0, asurface ¼ 1. The maximum absorbed heat flux is predicted by the analytical model to occur at the start of fire when the equipment is at ambient temperature. A similar set of values was used by the author to predict the front wall temperature versus time with the exception that the fire and gas temperature was set at 1472  F (800  C). Other combinations of values in the analytical model can be used that may provide an equal or better fit to the test data. Note that, because the vessel was engulfed in the pool fire, the gas temperature should be set equal to the fire temperature. The gas temperature will be lower than the fire temperature for non-engulfing pool fires. Also, the fire and gas temperatures were assumed to be constant throughout the pool fire. A comparison of the analytical model with the wall temperatures recorded during the test is shown in Fig. 4. The analytical method provides a reasonable approximation to the observed rear wall temperature versus time. The leveling off of the front wall temperature at about 800  F (425  C) as observed in the test (see Fig. 3b) cannot be approximated with a single set of parameters, indicating that one or more parameters changed during the course of the fire. Once the temperature versus time profile is approximated by the analytical method, then the resultant vessel heat input (i.e., qabsorbed) can be determined. The analytical method will indicate the maximum heat input is at the start of the pool fire where the vessel wall temperature is the lowest unless parameters change during the course of the fire or there was an instrument fault. 4.2. Use of the analytical method to model the Federal Institute for Materials Research and Testing (BAM) pool fire test (Balke, Heller, Konersmann, & Ludwig, 1999; Ludwig & Heller, 1999) wall temperature versus time In 1999, a full scale pool fire test was performed by the Federal Institute for Materials Research and Testing (BAM or Bundesanstalt für Materialforschung und -prüfung) in Germany (Balke et al., 1999; Ludwig & Heller, 1999). The test evaluated and compared fire exposure effects on a rail tank car containing propane and a Castor car used to transport radioactive material. The test setup is shown

Fig. 4. Comparison of rail tank car wall temperature versus time between the analytical model and Ballistics Research Laboratory pool fire test data.

Fig. 5. BAM pool fire test setup involving propane rail tank car (Balke et al., 1999; Ludwig & Heller, 1999).

 Embankment dimensions: 60  50  6 m (197  164  20 ft)  Tank car capacity ¼ 12,000 gallons (45.36 m3) TeT length ¼ 5.95 m; (19.5 ft) Diameter ¼ 2.9 m (9.5 ft)  Tank car test pressure ¼ 28 bar (406 psia)  Tank car contained liquid propane  Fuel oil pool fire in troughs under the tank car and Castor container  Castor container is used to store and transport radioactive materials and was tested along with the tank car. in Figs. 5 and 6. Although the tank car was semi-confined by embankments on 3 sides, a light to calm northerly wind was still able to significantly affect the pool fire exposure of the tank car as shown in Fig. 7. The tank car maximum pressure reached 25 bar (362 psig) about 15 min after the start of the pool fire at which time the tank car ruptured, resulting in a boiling liquid expanding vapor explosion (BLEVE). The BLEVE aftermath is shown in Fig. 8. Pool fire flame/gas and tank car wall temperatures versus time at the various locations around the tank car, as shown in Fig. 9, are illustrated in Figs. 10 and 11, respectively. It is important to note that failure occurred before the pressure reached the pressure relief device opening pressure. A discussion of this failure as it relates to depressuring system design is given in Section 6.3. Because of the wide pool fire and wall temperature ranges shown in Figs. 10 and 11, a single set of values for the parameters in the analytical method would not predict all variations. Only temperature data was given in the report. As in the BRL comparisons,

Fig. 6. BAM pool fire test setup involving propane rail tank car (Balke et al., 1999; Ludwig & Heller, 1999).

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test data. The rail tank car was filled with about 2650 gallons (10 m3) of 95% liquid propane, resulting in an initial wetted surface area of about 249 ft2 (23.16 m2). The pool fire heat input determined by the author for the empirical method without adequate drainage (see Equation (2)), and the analytical method for several locations around the tank car are given in Table 4. There are two locations in the rear of the tank car where the analytical method indicated higher heat inputs than the empirical method. However, when averaged across the entire tank car, as one should do if sizing a pressure relief device, the empirical method resulted in about 30% more heat input than the analytical method. 5.2. Comparison with pool fire heat input based on BAM fire test liquid sensible heating

Fig. 7. BAM pool fire test (Balke et al., 1999; Ludwig & Heller, 1999). Note: taken near end of test (calm wind speed).

unspecified parameters had to be determined by the author by trial-and-error whereby values were selected so the calculated time-versus-temperature profile matched the fire test data as close as possible. Table 3 illustrates the values of parameters selected to model tank car wall temperature versus time at two locations. Other combinations of values in the analytical model can be used that may provide an equal or better fit to the test data. It should be noted that a transient approach to the analytical method, where the fire and gas temperatures were varied with time based on the test data shown in Fig. 10, was evaluated; however, did not appear to significantly improve the fit with the test data. A comparison of the analytical model with the wall temperatures recorded during the test, shown in Fig. 12, indicates the analytical method can provide reasonable approximations to wall temperatures versus time if parameter values are empirically determined by selecting those that would provide a temperaturetime profile similar to the test data. 5. Comparison of the pool fire heat inputs between the empirical method, the analytical method and pool fire test data 5.1. Comparison with pool fire heat input based on BAM timeversus-temperature test data The pool fire heat input determined by the empirical method and the analytical method can be compared using the BAM pool fire

Fig. 8. BAM pool fire test e BLEVE aftermath (Balke et al., 1999; Ludwig & Heller, 1999).

Test data on the sensible heating of the propane liquid was obtained during the BAM test. This data can be used as an independent means to determine pool fire heat input during the BAM test. Note that the pressure did not reach the pressure relief device opening pressure prior to rail tank car failure during the BAM test. The test data indicated an average temperature rise of 8.06  F/min (4.48  C/min). Hence, the calculated total heat input due to sensible heating of the liquid is about 3.805  106 BTU/h (1115 kW). For comparison, the empirical method (assuming inadequate drainage) predicted a total heat input of 3.18  106 BTU/h (933 kW) per Table 4. This is roughly the same as the analytical approach using only the averaged tank car rear temperature data. A possible reason for these differences is discussed below. Liquid swelling as the liquid heats up would increase wetted surface area; however, the temperature did not increase enough during the test for it to explain the difference between the test data and the empirical and analytical methods. A likely explanation is that the embankment on three sides of the tank car heated up during the fire and caused higher heat fluxes due to re-radiation, preheating of combustion air, and enhanced heat transfer. Indeed, the fire should be classified as semi-confined because the height of the embankment walls exceeded the height of the tank car. In such cases, the API 521 empirical method (Equations (1) and (2)) does not directly apply. However, the equations can be modified by using a wetted surface area (Aw) exponent of 1.0 instead of 0.82. This would be appropriate in scenarios where the pool fire flames directly and continuously contact all of the wetted surfaces. Applying this to the BAM test rail tank car assuming 50% of the rail tank car is partially confined due to the embankments on three sides results in the Equation (4):

Fig. 9. BAM pool fire test e Temperatures measurement locations (Balke et al., 1999; Ludwig & Heller, 1999).

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situations (e.g., enclosed buildings or structures with a roof) which would require special modeling. Note that the empirical method is based on full-scale test data, not theory. Hence, this adjustment of the wetted exponent for partially confined vessels based on test data is consistent with the origin of the empirical method. 5.3. Comparison with pool fire heat input based on the BRL test

Fig. 10. BAM pool fire test e Fire temperatures versus time (Balke et al., 1999; Ludwig & Heller, 1999). Note 1: Time ¼ 0 is when gasoline starter fluid in a small plastic container was ignited. The main pool fire started about 100 s later when the plastic container failed and spilled burning gasoline into the fuel oil pool. Note 2: The unmarked temperature curves were primarily in the front of the tank car (upwind location and without an adjacent embankment).

Q API modified empirical method i h ¼ 70; 900* ðAw confinedÞ1:0 þ ðAw openÞ0:82

(4)

h i Q API ¼ 70; 900* ð11:58Þ1:0 þ ð11:58Þ0:82 ¼ 1:349  106 Watts ¼ 4:604  106 BTU=h This is a conservative estimate of the total heat input as compared with the 3.805  106 BTU/h (1115 kW) determined from liquid sensible heating. Based on the test data, the analytical method should use the rear averaged heat input predicted by the analytical model (i.e., 3.75  106 BTU/h (1098 kW)) to obtain a reasonable approximation. Where validating data is unavailable, the highest heat input obtained from the analytical model should be used. The fire relief load can be determined by dividing the fire heat input by the heat of vaporization of the fluid at relieving pressure. These results indicate that the API empirical method can be applied to some semi-confined configurations, where adjacent embankments exceed the vessel height, by using a wetted area exponent of 1.0 instead of 0.82 for the portion of the vessel adjacent to the embankment. This would not apply to completely confined

The BRL test obtained data on the relieving rate versus time, which was compared with that obtained with the empirical and analytical methods. A transient approach was used by the author in these methods whereby the relief rate was varied with time to correspond to the decrease in wetted area as fluid is relieved. A comparison of the actual relief rate and that predicted by the empirical and analytical methods is given in Fig.13. Both the empirical and analytical methods predicted a decrease in relief rate with time because the wetted area is decreasing as fluid is relieved. However, the test indicated the relief rate actually increased with time. One explanation that can increase the relief rate versus time is that there was an increase in heat flux with time due to heating of the surroundings. Adjustments to the analytical method were made to account for enhanced heat transfer due to heat-up of the surrounding embankments during the test. Fig. 13 illustrates a modified analytical method where the convective heat transfer coefficient and the vessel absorptivity were increased by 20% every 2.5 min with a limit of 1.0 for the absorptivity. These were empirically determined by trial-and-error so predicted relief rate versus time matched the fire test data as close as possible. This should be considered an example of adjustments that can be made to adjust the model to fit test data, but they may not represent actual conditions nor be applicable to other fires. Note that the fire and gas temperature were not adjusted because the data showed the fire temperature to slightly decrease during the test. Instead of modifying the analytical method, the empirical method can be adjusted to account for the apparent increased heat input with time by increasing the exponent on the wetted area versus time. This effect can be illustrated by inserting the heat input determined from the actual relief rate and the wetted area in the empirical method for inadequate drainage; the equation is then solved for the wetted area exponent versus time. The results, shown in Fig. 14, indicate the wetted area exponent approaches 1.0 toward the end of the test. This suggests that flame contact with the entire vessel surfaces increases with time. Using a wetted area exponent of 1.0 for the entire vessel (located in a pit with embankments exceeding the vessel height on all sides), with the empirical method, would provide a conservative pressure relief system design. 5.4. Application to pressure relief device sizing

Fig. 11. BAM pool fire test e Tank car wall temperatures versus time (Balke et al., 1999; Ludwig & Heller, 1999).

The analytical model can be used to estimate the time-versustemperature profiles of unwetted walls of vessel exposed to pool fires. API 521 provides recommended values for the parameters in this model that should provide a reasonable approximation for a wide range of pool fires. However, the comparisons with full-scale fire tests indicate the model parameters would need empirical adjustment based on actual full-scale test data in order to more closely predict a specific pool fire behavior. Unfortunately, the state-of-the art is insufficient to allow a validated theoretical basis for adjustment that would apply to pool fires in general. Hence, the user is cautioned when applying the analytical model to pressure relief device sizing for the fire exposure scenario because improper selection of the parameter values can result in underpredicting the fire heat input to a vessel and consequent undersizing of the pressure relief device. Where validating fire test data is unavailable,

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Table 3 Analytical method parameters used to model tank car wall temperature versus time at two locations of the BAM pool fire test tank car. Parameter

Rear center

Fire temperature,  F ( C) 1832 (1000) Gas temperature,  F ( C) 1832 (1000) 3.52 (20) Convective heat transfer 2  2 Coefficient, BTU/h ft R (W/m K) Fire emissivity 0.6 Metal emissivity 0.5 Metal absorptivity 0.5 The following were calculated using the values above: Calculated initial incident heat flux, 34,570 (109.0) BTU/h ft2 (kW/m2) 20,321 (64.1) Calculated maximum absorbed heat flux, BTU/h ft2 (kW/m2)

Front right 482 (250) 482 (250) 3.52 (20) 0.6 0.4 0.4 2260 (7.1) 1721 (5.4)

the user should use the API 521 empirical method for sizing pressure relief devices for the pool fire scenario. The comparisons in the previous sections indicate the 50 þ year old empirical method in API 521 given in Equations (1) and (2) would provide a reasonable approximation of unconfined pool fire heat input to a vessel but would need adjustment for cases where the vessel is partially confined. If the vessel is partially confined, the BRL full-scale test indicates a conservative size for a pressure relief device can be obtained using the heat input from Equation (2) but with a wetted area exponent of 1.0 instead of the 0.82 exponent valid for unconfined pool fires as shown in Equation (4). API 521 does not provide any recommendations regarding completely confined pool fires. In the case of completely confined pool fires, there is no full scale test data available so it cannot be determined if the API 521 empirical methods would be conservative. Although the analytical method may provide the potential to model completely confined pool fires, selection of the values for the model parameters would need to consider their variation with time as the fire heats the surroundings. This can only be done on a case-by-case basis. Further, validation tests to support selection of appropriate values for the model parameters would be recommended. 6. Application of the analytical method to depressuring system design 6.1. Effect of overheating unwetted metal plates Unwetted metal surfaces are not cooled by boiling liquid inside the vessel. Hence, the metal temperature can get high enough such

that metals such as carbon steel lose significant strength. Table 5 illustrates the effect of high temperatures on the tensile strength of carbon steel and 304 stainless steel. The loss of strength due to pool or jet fire exposure could exceed the safety factor used in the design of the vessels, thereby resulting in vessel rupture due to overheating, rather than overpressure. The specific pressure vessel design code and material used will determine the appropriate safety factor to use in the vessel design. For example, the current edition of ASME Section VIII, Division 1 Pressure Vessel Design Code (ASME Section VIII) includes a safety factor (now termed “design margin”) of 3.5 between the tensile strength of the vessel and the allowable stress at room temperature for materials in which the tensile strength governs (e.g., carbon steel). For carbon steel, the safety factor implies the design pressure is a minimum of 3.5 times the burst pressure (assuming the weak link in the vessel is the wall plate, there are no imperfections in the wall, etc.). It should be noted that carbon steel vessels constructed to pre-1999 versions of the ASME Section VIII, Div. 1 code used a safety factor of 4. ASME Section VIII, Division 1, UG-27 provides equations that relate the allowable stress, vessel design pressure and wall thickness. In the case of circumferential stress for a cylindrical shell, Equation (5) applies in many cases:

P ¼ S*E*t=½R þ ð0:6*tÞ

(5)

where: P must be < 0.385 *S*E. P ¼ internal design pressure, psi (MPa) E ¼ joint efficiency ¼ 1.0 for full X-ray. S ¼ maximum allowable stress value, psi (MPa) t ¼ minimum thickness of the shell, inches (mm) R ¼ inside radius of the shell, inches (mm) For example, at room temperature, ASTM A515 Grade 70 carbon steel plate has a tensile strength of 70,000 psi (482.6 MPa) (ASME Section II, 2007) therefore, the allowable stress will be S ¼ 70,000/3.5 ¼ 20,000 psi (137.9 MPa). If a vessel fabricated from this material and designed to this allowable stress is heated to 1200  F (650  C), the tensile strength will decrease to 20,000 psi (137.9 MPa), as shown in Table 5. In other words, the material strength is reduced to the equivalent of a zero safety factor. Vessel rupture would be a certainty if the pressure then exceeded the design pressure because the loads on the vessel would exceed the tensile strength. Rupture would occur at even lower internal pressures if there are other coincidental loadings on the vessel (such as the weight of the vessel and attached equipment, temperature gradients, static head, internals, etc.) or defects in the vessel. In all these cases, a pressure relief valve would not provide protection because it is typically designed to reseat (i.e., close) when the pressure decreases to about 93% of its set pressure for vapor trim valves. This would thereby maintain the vessel pressure near its design pressure. Instead of a pressure relief valve, a depressuring system can be used to provide vessel protection, or at least mitigation of the effects of failure. 6.2. Depressuring criteria

Fig. 12. Comparison of rail tank car wall temperature versus time between the analytical model and BAM test data.

In order to be effective, the depressuring system needs to depressure at a high enough rate to compensate for the loss of strength as the vessel heats up. The vessel heat up rate is dependent on the type of fire, materials of construction, and wall thickness. API Std. 521 Fig. 1 illustrates the heat-up of carbon steel plates of several thicknesses in an open pool fire (ANSI/API Standard 521,

E. Zamejc / Journal of Loss Prevention in the Process Industries 27 (2014) 21e31

29

Table 4 Pool fire heat inputs using the empirical and analytical method along with BAM fire test data. Location

Max fire heat flux, BTU/ft2 h (kW/m2)

Max absorbed heat flux, BTU/ft2 h (kW/m2)

Aw exponent

Total heat input, BTU/h (kW)

Analytical e Rear center Analytical e Front right Analytical e Rear average (Note 1) Analytical e Front average (Note 1) Analytical e Total average Empirical method

34,560 (109) 2260 (7.1) 24,350 (76.79) 6650 (20.97) 15,500 (48.88) N/A

20,330 (64.1) 1720 (5.4) 15,040 (47.41) 4620 (14.57) 9830 (30.99) 34,500 (70.9) (Note 2)

1 1 1 1 1 0.82

5.07 0.43 3.75 1.15 2.45 3.18

     

106 106 106 106 106 106

(1485) (126) (1098) (337) (718) (933)

% Of API 159% 14% 118% 36% 77% 100%

Note 1: Average of left, center and right locations. Note 2: The API maximum absorbed heat flux has units of BTU/h/[(ft2)0.82] or kW/[(m2)0.82].

2013). One curve (Plate 2) was obtained from open pool fire test data while the others were extrapolated based on the test data. Combining these temperature-versus-time curves along with the tensile strength data shown in Table 5 will allow determination of a minimum depressuring rate to keep the pressure below the tensile strength of the vessel. An appropriate safety factor should be considered given the uncertainties. Results obtained by the author, applying a 25% safety factor (i.e., Table 5 tensile strengths were multiplied by 0.75), are shown in Fig. 15. The depressuring profile for a specific wall thickness needs to stay to the left of the specific curve shown in Fig. 15. Failure will occur if the depressuring profile either intersects or is on the right side of the curve for the thickness in question. As noted in the previous section, failure can occur at even lower pressures, depending upon the amount of additional loads on the vessel. An often used criteria for depressuring is to depressure to 50% of the design pressure in 15 min. As shown in Fig. 15, this would be appropriate for open pool fire exposure of vessels whose wall thickness is 1 inch or greater. A second criteria often used is to depressure to 100 psig (6.90 barg) in 15 min. This is generally more conservative given the high pressures involved in most depressuring applications. The more stringent criteria, (almost always the latter) would be preferred when protecting against jet fire exposure.

Fig. 13. Empirical and analytical method calculated relief rates versus BRL test data.

6.3. Application of the analytical method to depressuring system design The analytical method can be used to extend the curves in Fig. 15 to other wall thicknesses. The analytical method along with the parameters determined in Section 6.2, for example, can be set up in a spreadsheet as a transient model in which the wall temperature change with time is calculated. At each time interval, the metal wall mass can be conservatively assumed to absorb all of the heat input, thereby increasing the wall temperature. The effect of wall thickness is accounted by the metal mass. This temperature-versus time profile is then combined with tensile strength data as in Section 6.2. For example, the BAM pool fire test data indicated failure of the rail tank car occurred at rear center wall (in unwetted zone) (Balke et al., 1999; Ludwig & Heller, 1999). Test data further indicated the wall temperature ranged from 1020 to 1200  F (550e650  C), but it is possible that local temperatures got even higher because temperature was recorded only at a few locations. Failure occurred 15 min after the start of the pool fire, or about 10 min after the fire temperature reached about 1832  F (1000  C). The rail tank car wall thickness was 0.59 inches (14.9 mm) and the material of construction was assumed to be comparable to carbon steel. The failure pressure of 362 psig (25 bar) was slightly lower than the test pressure of 406 psig (28 bar). The “Rear Center” parameters were used in the analytical model to predict the time-versustemperature profile. This was combined with the tensile strength and stress rupture data by the author to obtain the depressuring profile shown in Fig. 16. In order to minimize the potential for rupture due to overheating, a depressuring system would need to stay to the left of the curve shown in Fig. 16. Because the pressure at failure was slightly lower than the test pressure, Fig. 16 predicts that

Fig. 14. Empirical method wetted area exponent versus time using BRL test data.

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E. Zamejc / Journal of Loss Prevention in the Process Industries 27 (2014) 21e31

Table 5 High temperature tensile strength of carbon steel and 18-8 stainless steel (Wharton, 1946). Temperature F



900 1000 1100 1200 1300 1400 1500 1600

Temperature C



482 538 593 649 704 760 816 871

18-8 Stainless steel (304, 304L) Tensile strength psi

Tensile strength MPa

53,000 48,500 43,000 35,000 27,000 20,500 17,650

365.4 334.4 296.5 241.3 186.2 141.3 121.7

Carbon steel (SA-515, SA-516) Tensile strength psi

Tensile strength MPa

45,500 36,500 27,200 20,000 13,500 9025

313.7 251.7 187.5 137.9 93.1 62.2

Fig. 16. Depressuring profile to minimize failure potential of the rail tank car due to overheating in the BAM pool fire test.

failure would occur about 14 min after the start of the pool fire, which is a reasonable approximation as failure actually occurred about 15 min after the main pool fire started (see Figs. 10 and 11). Note the first 2 min of the pool fire test is not considered because the fire was localized to a small igniter assembly that did not cause any significant increase in rail tank temperatures.

The increasingly widespread use of analytical methods to evaluate fire exposure of equipment prompted the API Std. 521

committee to include an analytical method in the 6th edition as an alternative to the existing empirical method. The analytical method provides more flexibility than the empirical method but has limitations (e.g., too many permutations are possible). API 521 provides recommended values that would apply to many open pool fires. However, the comparisons with full-scale pool fire tests indicate that caution needs to be taken when selecting the values; otherwise, the fire heat input can be underestimated resulting in an undersized pressure relief device. Where uncertain, the values selected in the analytical model should be validated with testing. More recent pool fire test data indicates the 50 þ year old API 521 empirical method will provide a conservative estimate of pool fire heat input for most applications and is still the method of choice when designing pressure relief systems for an open pool fire scenario. However, these recent tests indicate the empirical method needs to be modified when a vessel or equipment is partially confined by adjacent embankments or walls equal or greater than the vessel height. In such cases, the wetted area exponent should be 1.0 instead of 0.82. The analytical method is useful in determining time-versustemperature profiles for heating unwetted vessels of varying wall thicknesses and materials of construction. These profiles can be combined with tensile strength and stress-rupture data to specify a depressuring system’s pressure-versus-time profile to minimize failure and/or mitigate the effects of failure due to overheating from a pool or jet fire exposure.

Fig. 15. Reduction of carbon steel plate tensile strength versus time due to open pool fire exposure.

Fig. 17. Depressuring profiles to minimize failure potential of a 0.5 inch wall thickness carbon steel and stainless steel vessel due to overheating.

6.4. Effect of material of construction The material of construction can significantly affect the depressuring requirements. The preceding sections discussed carbon steel vessels. As shown in Table 5, 304 stainless steel is superior to carbon steel regarding high temperature effects on tensile strength. A comparison of the depressuring profiles to minimize the potential for failure of a ½ inch wall thickness carbon steel vessel and a ½ inch wall thickness stainless steel vessel is illustrated in Fig. 17. The depressuring system pressure-versus time profile would need to stay to the left of the applicable curve. These results indicate that the depressuring system for the stainless steel vessel would require a significantly lower depressuring rate than for the carbon steel vessel of comparable wall thickness. This method can be extended to other materials provided tensile strength data at high temperature is available. 7. Conclusions

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References Anderson, C., Townsend, W., Zook, J., & Cowgill, G. (September 1974). The effects of a fire environment on a rail tank car filled with LPG. FRA-OR&D Report Number 75e 31, PB-241358. ANSI/API Standard 2000. (November 2009). Venting atmospheric and low-pressure storage tanks (6th ed.). ANSI/API Standard 521. (2013). Pressure-relieving and depressuring systems (6th ed.). API Standard 650. (2013). Welded tanks for oil storage (12th ed.). ASME Section II, Part D, materials e Properties. (2007). ASME Section VIII, Division 1, Pressure Vessel Code, 2007 with 2008a Addenda. Balke, C., Heller, W., Konersmann, R., & Ludwig, J. (September 13, 1999). Study of the failure limits of a railway tank car filled with liquefied petroleum gas subjected to an open pool fire test. BAM Final Report. Energy Institute. (March 2003). Guidelines for the design and protection of pressure systems to withstand severe fires, ISBN 0 85293 279 0. Heller, F. J. (1983). Safety relief valve sizing: API versus CGA requirements plus a new concept for tank cars. API Refining Proceedings, 82, 123e140. Ludwig, J., & Heller, W. (1999). Fire test with a propane tank car. BAM Test Report III.2/9907.

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National Fire Protection Association NFPA 30 (2008): Flammable and Combustible Liquids Code. Personal correspondence from H.C. Hottel to L.W.T. Cummings December 12, 1950. Roberts, T. A., Medonos, S., & Shirvill, L. C. (June 2000). Review of the response of pressurised process vessels and equipment to fire attack. Offshore Technology report, OTO 2000-051. Salater, P. Proposed changes to the next revision of API 521. 2006 Presentation to API Pressure Relief Systems Committee. Salater, P., & Overa, S. J. (March 2004). Pipes exposed to medium sized jet fires e Rupture conditions and models for predicting time to rupture. Paper presented at Fire and Blast Information Group (FABIG), London and Aberdeen, January 2004 and Houston. Salater, P., Overa, S. J., & Kjensjord, E. (September 2002). Size depressurization and relief devices for pressurized segments exposed to fire. Chemical Engineering Progress, 38. SCANDPOWER. (March 31, 2004). Guidelines for the protection of pressurised systems exposed to fire. Report no. 27.207.291/R1-Version 2. Shirvill, L. C. Heat Fluxes in Severe Fires. 2002 Presentation to API Pressure Relief Systems Committee. Wharton, H. R. (1946). Digest of steels for high-temperature service. Timken Steel.