An approach to hazard analysis of LNG spills

Journal of Occupational Accidents, 7 (1986) 251-272 Elsevier Science Publishers B.V., Amsterdam - Printed in The Netherlands

AN APPROACH

TO HAZARD

ANALYSIS

251

OF LNG SPILLS*

D.H. NAPIER and D.R. ROOPCHAND Deparfmenf of Chemical Engineering and Applied Chemistry, Toronto, Ontario, M5S lA4 (Canada)

University of Toronto,

(Received 30 September 1985; accepted 7 November 1985)

ABSTRACT Napier, D.H. and Roopchand, D.R., 1986. An approach to hazard analysis of LNG spills. Journal of Occupational Accidents, 7: 251-272. The extensive demand for fuel gas in modern communities has led to the construction of liquefied natural gas (LNG) terminals and peak-shaving facilities; the latter are often in close proximity to urban areas. Such installations carry with them actual and other perceived hazards; the analysis and control of these hazards are important. In this paper a logical approach to hazard analysis is described and details of the features of the analysis are outlined. A commentary on dispersion models leads to recommendations for the choice of a suitable dispersion model. Combustion and explosion are also considered and limiting features of acceptability are discussed. Part of this discussion is comprised of an evaluation of the adequacy of thermal radiation models from pool fires. Three methods for predicting radiant flux densities that would be received at the property line are examined and compared to those deduced from the Canadian Standard 2276 which is similar to the NFPA Standard 59A. The basis for setting acceptable minimum values for thermal radiation for both pilot and spontaneous ignition of cellulosic materials is also discussed. In conclusion, various aspects of hazards arising from LNG are considered and broad guidelines for analysis are set down.

INTRODUCTION

The excellent fuel properties of natural gas, combined with both its value as a petrochemical feedstock and its relative abundance, has led to its increased use in urban domestic and industrial environments. Natural gas is often not found in areas where it is in great demand, hence the need for extensive transportation networks and distribution centres. To facilitate transportation, the gas is liquefied under refrigerated conditions to form a clear, odourless liquid with a boiling point of -162°C and a density of 470 kg/m’. The liquefaction process involves a volume reduction factor of about 600 from the gaseous state. *Presented in part at the 33rd Canadian Society Toronto, Ontario, 1983.

0376-63~9l861$03.50

for Chemical Engineering Conference,

o 1986 Elsevier Science Publishers B.V.

252

An investigation into the hazardous properties of liquefied natural gas (LNG) is appropriate as Canada is on the verge of embarking upon very large LNG projects on both of its coasts. The aims of this paper are therefore: (a) to investigate some of the hazardous properties of LNG and (b) to present the results in a manner suited to hazard analysis. The goal of hazard analysis is to quantify risk. Once quantified, the degree of risk forms a basis for: 1. Comparison with other acceptable risks 2. Assessing improvements in plant safety that can be gained by modifications in the design stage. 3. Siting plant and pieces of plant relative to other plants (or pieces of plant) or residential areas 4. Assessing industrial insurance premiums 5. Various aspects of urban planning Within this paper LNG is examined in terms of release, dispersion, combustion and rapid phase transition (RPT) phenomena. HAZARD

ANALYSIS

The four general steps in performing a hazard analysis are [l] : identification of failure modes, calculation of consequences, estimation of failure probabilities, and assessment of the overall impact. At the onset an inventory of the hazardous materials is prepared to ascertain the quantities, temperatures, pressures and locations of hazardous materials. Failure scenarios are then postulated by carefully examining the Process Flow, and the Process Piping and Instrumentation diagrams, as well as previous accident reports where applicable. There are several established methods available in the form of checklists [2] and hazard and operability studies [ 3, 41 to simplify this exercise. In essence, the study progresses from calculations of the discharge, the boil-off and spreading rates, to modelling both the dispersion of the vapour cloud using heavy gas dispersion models, and the combustion phenomena which may ensue. The study then concludes with a summary of the overall impact of all possible failure scenarios. The range of impacts will ultimately depend upon the prevailing meteorological conditions, the local terrain, and the relative locations of ignition sources, buildings, and persons to the spill site. RELEASE

The final outcome of a release of LNG will depend on the following release characteristics and may, in general, range from a localised fire to a completely non-combustive scenario where the primary hazards are those of asphyxiation and extreme cold: l the physical properties of the flammable liquid l the temperature and pressure conditions prior to loss of containment l the quantity, type and duration of release

253

0 point of release time delay to ignition. The degree of pressurization prior to release will be collectively determined by the physical properties of the liquid and the storage conditions. Many of the failure scenarios postulated will result in non-pressurized releases where LNG spills upon a surface and rapidly vaporizes. In those few scenarios where severe internal overpressurization is postulated, momentum jets with two-phase flow and rapid vapour flashing with fluid entrainment are possible. The scope of the problem is defined by the quantity of material released, which in turn is dependent on both the size of the containment system and that of the leak. In general, LNG facilities, whether they are export, import or storage terminals, consist of one or more of the following: a jetty to handle ships or a loading area for railcars and trucks, a transfer system, a pressure control system to control boiling-off produced in storage tanks and vapour balancing during transfer operations, a liquefaction facility (often with flammable refrigerants), and lastly a high pressure gas dispatch system usually with a vaporization facility. In the dock and shipping areas, potential releases may occur from loading arms, large diameter transfer lines, circulation lines, surge drums and, where applicable, the LNG carrier itself. Typical release rates and durations are provided in Tables 1 and 2, and are limited by emergency shut-down provisions. In the absence of such shut-down provisions, spill durations should be lengthened accordingly. These tables can be used to calculate the quantity of material that would be spilled in any of the failure scenarios listed. The failure rates can be used in hazard analyses to both quantify and compare risks. Risk is measured as the product of the consequences (that may be measured in terms of injuries, lives or material damage) of an event and the frequency of occurrence of that event (failure rate in eventslyr). For example, if it were necessary to compare the risks involved in accidents which resulted in either a severed loading arm, a severed 36-inch transfer line, or a break in the absorber bottom connection, the following approach can be taken. From Tables 1 and 2 the following data are extracted. l

Failure

Mass released (kg)

Frequency (events&r)

Loading arm Transfer line Absorber connection

13500 2400 X 60 X 5 min = 720,000 3000 x 19s = 57000

1 x 10-S 4 x lo-’ 1 x 1o-5

Risk, as mentioned previously, is the product of the consequence of the failure (in this case the mass of LNG released on which depends the resulting loss of life, injuries and material damage) and the frequency. The risks for the three failures are 0.135, 0.288 and 0.57 kg/yr, respectively. Therefore, there is more risk involved with the absorber bottom connection than with

254 TABLE

1

Typical

failure scenarios

[ 1]

-~Scenario

Storage tanks 1. Catastrophic failure of inner tank leading to outer roof collapse 2. Partial fracture of outer roof due to overpressurization 3. Catastrophic rupture of primary and secondary containment 4. Serious leakage from inner tank

Dock areas 1. Failure of loading arm together failure of the shutdown system 2. Failure of 36-inch transfer line Vaporisers 1. Absorber 2. Vaporiser

bottom connection manifold failure

with

failure

Refrigerant storage 1. Catastrophic failure of 480 tonne semi-refrigerated propane tank l/2 full 2. Catastrophic failure of refrigerated full ethane tank Slugcatcher 1. Total failure of slugcatcher 2. Total failure of slugcatcher typical operating conditions

under (10% full)

Liquefaction and fractionation units 1. Total failure of heavies removal column 2. Total failure of debutaniser column LNG carrier 1. Total failure of one cargo tank 2. Total failure of two cargo tanks

Total mass released (kg)

Typical failure frequency (events per yr)

depends on tank design and capacity depends on tank design and capacity depends on tank design and capacity depends on tank design and capacity

0.8-2 x lo-&

13500

2 x 1o-5

1

x 10-9

2 x 10-S

1

x lo-*

2400 kg/s for 5 min

4 x lo-’

3000 kg/s for 19 s 27 kg/s

1

240,000

x 10‘5 1 x lo-‘5

1 x 10-S

82000

2

x 10-6

1,800,OOO

1

x 10-S

180,000

13,700 7800

11,000,000 22,000,000

1 x 10-5

1 1

x 1o-5 x 1o-5

5 x 10-e 0.5 x lo--”

the other two lines. This joint would warrant more careful design or perhaps additional expenditure to ensure reliability. The LNG carrier presents the greatest hazard as the spill potential is several orders of magnitude larger than other postulated spills. Far inland spills the applicable Canadian Standard [5] offers guidance for predicting

255 TABLE 2 Failures considered in a typical baseload LNG plant f 34 ] Type of failure

LNG release rate

Release duration

Tank top line Liquefaction area pipe Liquefaction to tankage lines Loading pump suction line Loading pump discharge line Pump to dock lines Dock manifold area Ship transfer arm Refrigerant storage Fuel storage

Line capacity rate Unit production rate Total production rate Line capacity rate Line capacity rate Line capacity rate Nozzle/line discharge rate One arm rate Largest pipe capacity Largest pipe capacity

3.0 3.0 10.0 3.0 3.0 10.0 1.5 1.5 3.0 3.0

min min min min min min min min min min

the size of design spills. However, it does not provide guidance for predicting the size of spills due to catastrophic failure of a storage tank. LNG storage tanks are usually very strong, as they in general consist of a cryogenic inner tank, insulation, load bearing outer tanks of carbon steel and/or concrete, and beyond that a diked area capable of containing the full volume of LNG contained in the tank. Once the inner tank has failed it is unlikely that the carbon steel outer tank will be able to withstand the thermal shock, hence the need for bunded or diked areas. Tank designers have anticipated many of the stresses that can lead to tank failure, and the modern tank is vulnerable to perhaps only a direct aircraft strike, and a prolonged fire at close proximity. The pressure control system usually consists of low pressure compressors and a gas absorber or recondenser [ 11. Vapour line failures would result in minor releases whereas failures in the LNG feed lines to the absorber could lead to larger releases of flammable liquids. Liquefaction areas have a large potential for hazard as they involve a large number of material hold-ups in the form of reflux drums, columns, heat exchangers, etc. Also, there is often a large inventory of pressurized flammable refrigerants such as butane, ethane, ethylene and propane. Many of these refrigerants are capable of producing both Boiling Liquid Expanding Vapour Explosions (BLEVE) and Unconfined Vapour Cloud Explosions (UVCE) and are considered to present a greater hazard in storage than LNG. Liquefaction plants which are fed by pipeline and are required to perform gas separation may have slugcatchers and spherical receivers. These vessels are transiently operated at pressures of 60-100 bar and upon failure will result in large releases of gas and entrained condensate due to the rapid depressurization of the liquefied gases within. The principal area of concern in the gas dispatch system is the high pressure (80 bar) LNG feed pumps to the vaporisers. The hazard upon release

256

may be further increased as these areas are not required by the CSA Standard to have impounding areas to limit the spread of LNG. The type of release, i.e., gas phase, liquid phase, or both, will have a significant effect on the ensuing dispersion process. Condensed phases will in general lead to greater mass flow rates and will lead to the formation of a liquid pool. The dispersion process will also be slower as a result of the ratelimiting vaporization process. The duration of the spill may vary somewhere between the extremes of instantaneous and continuous releases. The instantaneous spills produce higher downwind concentrations than any other types of spills of equal volume. Increased period of release results in an elongation of the initi~ly circular cloud in the direction of the wind. The behaviour of the liquid spilled will be determined by those spill characteristics listed above and the rate of supply of heat from the surface. With LNG, the predominant source of heat is in the surface beneath the pool (ground or water). Upon release, the liquid will vaporize quickly as it contacts the surface; following this flash vaporization the liquid will spread and continue to vaporize at a lower rate. The spreading can be adequately modelled as a density intrusion [6] using the following equations: V=Vi+V~t-mlpL v = w2h dr/dt = (egAh)lh

(1) (21 (9

For spills on water n = (I- pL/pw), e = 1.6 [6] For spills on land A = 1, e = 2 The pool will continue spreading until it reaches a minimum pool thickness, which depends on the material upon which it is spilt (see Table 3) Upon reaching this minimum thickness (and maximum radius) the pool will break up into smaller sections. When LNG is spilled onto water, film boiling occurs. Ice formation does not occur except in those cases where the water volume is small, The heat flux (4) from the water, and hence the vaporization rate, remains relatively constant (0.196 kgjm2s) [7] and depends only on the difference between the ambient water temperature and the boiling point of the liquefied natural gas. 4 = 600 (T, - !&,) TABLE Minimum

W/m2

(4)

3 thickness

for pool spread

Surface

~in~murn

Rough, sandy soil Farmland, pasture Smooth sand, gravel Concrete, stone Calm water

25 20 10 5 2.8

pool thickness

{mm)

257

The evaporation

rate is given by:

dmldt = q/L = 600(T,

kg/m*s

- Tb)/L

(5)

Once the pool has started to break up, an empirical relationship, proposed by Feldbauer et al. [7], can be used to estimate the evaporation rate: M = M,exp

((-0.04)(

t, - tm)/pHm)

(6)

The evaporation rate for a spill on land is much lower than that for water and changes with time. The important parameters of the soil are thermal diffusivity, density, heat capacity and thermal conductivity (see Table 4). TABLE Thermal

4 properties

of some soils and concrete

Material

Concrete Soil (average) Soil (sandy, dry) Soil (moist, 8% water,

sandy)

[ 61

Density

Specific heat

Thermal conduct.

Thermal diffus.

(kg/m’)

(J/kg-K)

(w/m*K)

(m’/s-

2300 2500 1650 1750

961.4 836.0 794.2 1003.2

0.92 0.96 0.26 0.59

4.16 4.59 1.98 3.36

The heat flux to the liquid pool can be estimated q = C’k(T,-

using:

Tb)/(mt)%

(7)

C’ = 1 for impenetrable surfaces C’ = 3 for other surfaces. For both land and water spills, Eq. (8) can be used to estimate ration from a spreading pool: dm -= dt

Ck(Ta

-

L(7rcY)”

10 ‘)

’

Tb) s

0

2nr’dr’ (t-

t’p

the evapo-

63)

DISPERSION

The dense gas evolved will spread radially away from the spill site under the influence of gravity. The spreading velocity will be greater than that predicted by a Gaussian-type model. Under the influence of gravity the cloud will move upwind. The velocity of the cloud upwind will be less than the corresponding velocity downwind thereby resulting in an asymmetricallydistributed cloud around the spill site. With the entrainment of air, the temperature of the cloud is increased and the density decreased. Initially the cloud will be 35% denser than ambient air; if the cloud remained isolated

258

and exposed to warming influences (e.g., sun, ground) it would become 55% less dense than ambient air. During the warming-up it would lift off. In practice the dense gas is diluted with air, warmed thereby and its density is reduced. The dense cloud will, in general, be “pancake-shaped” and can attain a width to height ratio of 100. The effect of local terrain (buildings, local vegetation, topography and other obstacles) will cause increased dilution via wake effects. However, some local features, such as declivities and ditches, hinder dispersion and produce hazards that differ in nature from those of the dispersing cloud. In some cases, releases into the wake of a large obstacle may cause a hold-up, which will have the effect of converting an instantaneous release into one that is time-dependent. Similarly, local vegetation may cause increased dilution as well as a reduction in the advection speed of the cloud. Wake effects from the ruptured container should also be considered. Pertinent meteorological conditions include wind speed, wind profile (determined by surface roughness and atmospheric stability), temperature and humidity. The downwind distances to the LFL are increased in high winds due to the greater bulk movement of the cloud. Releases in stable atmospheres produce longer and wider LFL contours than those in unstable atmospheres, as stability tends to inhibit vertical mixing. Both temperature and humidity will affect the heat transfer rates to the cloud and the concentration at the edge of the visible fog produced by cloud. There are two general types of dispersion models available, the K-theory (eddy diffusivity) and the “top-hat” or “slab” models. In principle, the Ktheory models are superior as they attempt to predict the pertinent characteristics of the cloud by examining its fluid dynamic and thermodynamic properties and can readily be modified to incorporate new developments in the aerodynamics of cloud movement around obstacles. The 3-D version of these models superposes a cubical array of user-selected-sized cubes around the spill site. Mass, momentum and energy balances are written for each cube and solved numerically to account for fluid dynamic acceleration, fluid transport, and turbulent mixing for each time step after the spill. These models are also able to deal with obstacles by introducing no-flow cells into the grid. At present they cannot account for obstacles which would allow partial flow such as vegetation, fences, etc. K-theory models include ZEPHYR [8], MARIAH [ 93, SIGMET-N [lo], TRANSLOC [ll] , and DISCO [ 121. Slab models were originally Pasquill~ifford (P-G) models which were modified to account for gravitational effects and the inhibition of vertical mixing. In general, these models assume a cloud with a specific shape which entrains air across density interfaces at the cloud edges, and internal mixing occurs at a sufficiently rapid rate to render an approximately unform internal composition. Slab models involve a transition to P-G-type dispersion when the cloud is no longer negatively buoyant; some models incorporate an automatic transition, others require a manual transition. Top-hat models include those by Cox and Carpenter [ 131, Kaiser [ 141, Germeles and Drake [15],Fay [16],vanUlden [17] andpicknett [18].

259

Recently, there has been a new approach to dispersion modelling, whereby the two types of models are combined to form a “super-hybrid”. An example of such a model is HEGADAS [19] which assumes a specific shape in the form of similarity profiles and uses eddy diffusivities to account for transport phenomena. Table 5 provides information for selecting an appropriate dispersion model. Spill tests have been too small to establish conclusively the accuracy of these models. Based on limited experimental data, K-theory models were able to match actual spill data with the greatest accuracy. Second to those are the TABLE

5

An overview

of heavy gas dispersion

models

[231

Conditions

Models K-theory

Top-hat

type

ZEPHYR

SIGMET

MARIAH

DISCO

G&D

HEGADAS

Eidsvik

Fay

Y

Y

Y

Y

Y

Y

Y

Y

Y

Y

Y

Y

Y

Y

Y Y

Y

Y

Y

Y+

Y*

Y*

Y

Y

Y

Y

Y+

Y

Y+

n n n

Y Y Y Y

Y

Y

n

n

Y

n n

Y

Y

Y*

n

n

Y

Y

n

Y

n

n

Y

Y

n

Y

Y

n

Y

Y

Y

Y

Y n

n

n n

n

Y

Y Y

Y n n

Y

Y

n

n

Spills

Instantaneous Steady continuous point source line source confined area SOUFX

Variable cont. point line confined area spreading area Dispersing regimes open obstructed (wakes) topographical (var. terrain) Meteorology atmos. stability low wind speed humidity and heat of condensation Mechanistic features transition from heavy gas to P-G

y Y

Y n

Y

Y

Y

Y

Y(l)

Y

Y

Y

Y

Y

Y

n

Y

Y

Y

y

n

Y

”

Y

Y

Y

Y

Y

2

2

3

6

4

5

5

none

r -requires several repetitions. y - applicable. y* -line source perpendicular to wind direction. y+ -treated as a vertical rectangular source. n -not applicable. y(1) -in Gaussian portion of dispersion. 2 - not needed use rezoning. 3 -not needed automatic rezoning. 4 -transition occurs when cloud edge speed exceeds 5 -smooth and continuous. 6 -not needed.

wind speed.

260

advanced slab models (HEGADAS). The least accurate are the slab models. The true power of K-theory models and HEGADAS is realized in their ability to produce realistic LFL contours around the spill site. The hazard assessor should consider the following prior to choosing a model: (1) the accuracy required (2) the financial resources available (3) the applicability of the model (4) the availability of the model (5) the simplicity of the model. K-theory models, for example, are “state-of-the art” in terms of accuracy; however, they require large computing capacity, extensive input‘ data, considerable computer time (as accuracy increases with decreasing grid size) and are often proprietary. fn many circumstances the cost of K-theory simulations cannot be justified and the use of similarity models and slab models is strongly recommended. The latter tend to overpredict, thereby introducing an added margin of safety. Slab models continue to be used extensively in risk analyses [6] as they are inexpensive, readily available, i.e., not proprietary, easily programmable and require less input data, thus making them more “transportable”. Physical modelling has also been used to predict the behaviour of heavy gases. ModelZing involves the construction of a scale model of the site and the release of a heavy gas simulant in either wind tunnels or on water tables. This method provides good qualitative data and fair quanti~tive data regarding the effects of obstacles. The latter is due to their present inability to incorporate all the properties of the heavy gas simultaneously.

With respect to combustion phenomena, the primary goals are to: define acceptable thermal radiation doses based on both the thermal flux and the exposure time, and ascertain the principal modes of combustion available.

Thermal radiation poses a hazard to both persons and property, hence there is a need for two sets of aeceptahi~ity criteria. Compliance with the criteria for persons will in general provide an adequate safety margin for property. As a guideiine, human radiation susceptibility appears to be doserefated with few minimum intensity thresholds. The threshold for pain is thought to be about 4.7 kW/m’. Beyond that the following relationship exists between radiant flux and exposure time necessary to cause severe blistering of skin [ZO] : & = 5o/t0*71

(9)

261

Estimated relations between thermal radiation intensity and burn injury, based on nuclear explosion research by Glasstone and White, are found in Table 6 [Zl] . Criteria for property damage are often based on the flux and duration necessary to ignite cellulosic materials via both pilot and spontaneous ignition. The following equations (211 can be used to calculate exposure times at radiation intensities above the critical values listed below: for pilot ignition (I’ -

Ip)t2’3 = k,

(10)

where

lP is critical intensity for pilot ignition itt, is a constant 8050 J/m2 s1’3 for spontaneous ignition (I’ -

13400 W/m2

&)P” = h,

(11)

where

1, is 25400 W/m2 k, is 6370 J/m2 P Acceptability criteria can also be embodied in allowable radiation contours drawn around a potential fire site. Robertson [22] suggested the following contours: a 2.37 kW/m’ contour for continuous exposure of people fighting the fire, 4.70 kW/m2 for a 10 s exposure of personnel during the emergency, 12.6 kW/m’ for severe damage to property, cables and sensitive equipment (via ignition of cellulosic-type materials), and 37.8 kW/m2 for storage tanks. The CSA Standard also provides compliance equations to calculate separation distances from LNG impounding areas, (see Table 7). The compliance equations make provision for a flux of 30 kW/m2 at a property line which can be built upon. In comparison with other criteria, this appears to be too high as it is capable of igniting cellulosic material within 2 s. TABLE 6 Estimated relations between intensity and burn injury Injury

Radiation intensity (kW/m%) at various exposure times 1.5 s

10 s

45 s

73 131

12 27

4 9

73

12

4

Significant injury threshold

117

25

9

Lethality

143 264 586

33 58 128

10 19 40

First degree burn Second degree burn Slightly clothed, few if any injuries

threshold near 50% near 100%

262 TABLE

7

Canadian

standard

compliance

equations

Area of concern Point more

impounding

Acceptable @W/m’)

of outdoor assembly of 50 or persons beyond the property line

Nearest

for LNG

point

At a property upon

of a building line which

structure can be built

flux

areas Compliance

5

d’ = ,(A)“’

9

d’ = ,(A)]”

30

equation

(m)

d’ = 0.8(A)1’Z

Combustion In general, the principal modes of combustion from a spillage of highly volatile, combustible liquids are Boiling Liquid Expanding Vapour Explosion (BLEVE), Unconfined Vapour Cloud Explosion (UVCE), pool fire, cloud fire, confined explosion and detonation. Based on the physical properties of methane and various degrees of experimentation, LNG was not found to be a candidate for either BLEVE, UVCE or unconfined detonation phenomena [23, 241. Confined explosion, pool and cloud (flash) fires are the primary routes of combustion considered. Confined

explosion

Confined explosions may occur wherever flammable gas or liquid enters areas of confinement. Such explosions were responsible for much of the damage that took place after the release of approximately 4200 m3 of LNG in Cleveland, Ohio (1944). The damage potential of confined explosions can be estimated using traditional methods whereby the heat of combustion of the fuel is correlated with an equivalent mass of TNT. The TNT equivalent, and the distance to the target from the source of the explosion, can then be reduced to a scaled distance parameter. Plots of scaled distance versus overpressure are then used to estimate the amount of overpressure that will be experienced by the target. Tables are readily available to compare the extent of damage that may result from varying amounts of overpressure [21, 25, 261. Pool fire The development of either a pool or a cloud fire after an accidental spill depends on the pre-ignition period. Early ignition would result in the former and delayed ignition in the latter. To assess the hazards presented, it is necessary to examine the pertinent characteristics of the flame.

263

Flames above a pool of LNG differ from those over pools of other liquids in that they are more difficult to initiate and once ignited they produce much less soot, and they tend to spill over the confining wall of the pool; they have higher liquid regression rates, higher length to diameter ratios, and emit radiation largely in the carbon dioxide and water bands. The latter causes greater atmospheric absorption of radiation from an LNG fire than that from other hydroc~bon fires. The liquid regression rate for LNG pools on land is 2.35 X 10m4 m/s and for pools on water it is 4 X 1O-4 m/s [24]. The flame shape can be approximated to the shape of a cylinder with a diameter equal to that of the pool. The height of the flame can be estimated using the Thomas correlations [27] ; however, these correlations tend to underestimate experimental data [28] by about 13%. Several researchers have produced correlations for estimating both flame height, flame drag and flame tilt (29-311. They also recommend the use of tilted flame cylinders with elliptical cross-sections to better account for the flame shape variations due to spill-over and flame tilt. Nominal flame surface emissive powers are about 153 kW/m’ [28] with maximum values ranging to 270 kW/m’. The equivalent black body flame temperature is approximately 1150 K and the fraction of the total heat of combustion radiated amounts to 23%. Methods for estimating

thermal radiation fluxes

Three methods are presented for calculating the radiative flux from a hypothetical pool fire around an existing storage tank 31 m high and 42 m in diameter (storage capacity 41500 m3). The tank was bunded to provide a secondary level of confinement. The bund configuration and the target distance were chosen to be in compliance with the CSA Standard. The bund dimensions were 30 m high and 60 m in diameter and the target distance was the distance to the 5 kW/m2 contour (159 m from edge of bund to target). The emissive power of the flame and the flame temperature were assumed to be 153 kW/m2 and 1150 K, respectively. The results from the three methods of calcuation are presented in Table 8. The first method is used widely to calculate the incident heat radiation flux from flames on burning liquid pools and from flares [21]. The total TABLE 8 Predicted radiation intensity levels at the 5 kW/m contour for an LNG pool fire scenario Method

Radiation intensity at a distance of 189 m from centre of pool (kW/m’)

Method 1 Method 2 Method 3

6 9 4

264

Fig. 1. Source

flame representation

method

1 [21].

radiated heat from the flame is assumed to originate at a point source at the centre of the flame (see Fig. 1). The flame is assumed to be cylindrical in shape with a height equal to twice that of the pool diameter. The incident radiation intensity (I) at the target is given by: I = (Q’ cos 8)/4nR2

(12)

Q’ can be estimated as product of total flame surface area and emissive power. This technique is not considered to be applicable for target distances less than 3 pool diameters as relative geometry between the flame and the target become important (this is especially true during episodes of severe flame tilt and flame drag). The second method, proposed by Robertson [22], considers the flame to be a flat radiator with dimensions d X 2d (d is diameter of pool) and vertically disposed above the tank (see Fig. 2). The area of the radiator (a) viewed from the target location on the ground is (a’ ) such that a ’ = a co@. A very simple view-factor is calculated using the inverse square law, such that the radiation R, received at a distance x is approximated by: (13)

R, = &(e’f(nX:)) where 3c= (X, + ol sine)cose (14) and tane = (d + H)/x: (15).

Fig. 2. Flame

representation

for Method

2 [22 J.

265

The rate of heat radiation from the flame Rf is calculated with the assumption that the flame is a cylindrical surface radiator, hence Rf = Q’/27rdz. Although this method makes many assumptions with respect to relative geometries, it is widely used for separation of items of plant. The third method also assumes that the flame is a flat radiator with dimensions d X 2d and the appropriate configuration factors are calculated. The target is assumed to be parallel to the flat radiator to facilitate a maximal calculation. The configuration factor (I$) can be estimated from the following equations and is illustrated in Fig. 3. The $ is for an element of area placed in parallel with a rectangular surface where one corner of the rectangular surface intersects the normal from the centre of the area element. To satisfy the corner requirement in the definition of 4, the flat radiator is divided into equal sections A; and A;. Areas A; and Ai represent the nonradiating bund wall. Consider only areas A; and Ai and double the resulting r$ to give a result for the entire flame. @=

&[((l +BB*),,,)tan-’ ( +:2)li2)+ ( +cc2)l~2) tq(l +Bcy#l~ (1

B = b/f; C = c/f overall @ = ($1 - f#~~) X 2 @I for A,’ and Ai G2 for Ai The radiation intensity received I = @rue Tf

(1

at the target is given by: (17)

Atmospheric transmissivity and flame emissivity are also taken into consideration; however, they are not known accurately. Transmissivities are often well below unity due to the strong absorption by carbon dioxide and water; however, for this example it is assumed to be unity. An emissivity value of unity is used with a black body flame temperature of 1150 K [28]. In selecting a method of calculation the hazard assessor is advised to employ Method 3 as it provides the most accurate estimates; this is especially true when the view factor is calculated rigorously using summation methods over actual flame shapes. Excepting Method 3, the next most accurate and simplest to implement is Method 1, which is currently used in many Standards. Cloud (flash) fires Ground level cloud fires have not been subjected to the same intensity of research as pool fires; consequently there is much less information available. An LNG vapour cloud is not homogeneous and will burn initially in the peripheral regions of the cloud where the mixture is within the flammable limits. Once combustion has commenced, the remaining fuel-rich portions of the cloud will burn as a diffusion flame and will gradually move towards

266

CONFIGURATION FACTOR ($1

-RADIATING

AREAS

1

FL

Fig. 3. Flame

geometry

used in Method

3 [23 1.

the source of the vapour, i.e., the liquid pool. On arrival there, the flames will stabilize and the combustion will proceed as a pool fire. Cloud fires are of relatively short duration and do not present a large radiation hazard to plant; instead, they act primarily as a source of pilot ignition. Cloud fires rapidly reduce the oxygen concentration around the fire and hence pose an asphyxiation hazard to persons. The flames move with normal flame speeds, increased somewhat by the prevailing atmospheric turbulence. They do not accelerate towards the source of the vapour, nor do they exhibit any evidence of fireball-type behaviour. The last aspect to be considered is that of rapid phase transition explosions (RPT) [32]. These explosions arise from the contact between dissimilar liquids at different temperatures. RPT explosions have been observed during spills of LNG on water. The resulting explosions are mild in comparison to explosives such as TNT. The maximum theoretical explosion pressure and energy yields are 30 bar and 124 kJ/m*. RPT explosions present a mild

267

blast hazard. More importantly they have been shown [33] to present a potential source of ignition for the consequent flammable vapour cloud. Ignition may arise from two possible sources: (1) charge on droplets produced in the explosion (as in an LNG spillage on water) collecting on an isolated conductor (e.g., a slug of water) (2) friction sparks from impacts between metallic objects, that are brought about by the blast wave SUMMARY

OF THE OVERALL

IMPACT

A summary of the impact of all possible scenarios, which includes probabilities and possible consequences for each failure case in each of the possible weather conditions, must be made to produce credible risk statistics. These failure scenarios may number in the 1000’s for a single plant, hence the need for an organized approach and concise methods of summarization. The goal is to provide some measure of the total risk incurred in operating the facility. The societal and individual risk are expressed in terms of the number of probable deaths. Societal risk involves the possibility of an accident causing multiple casualties, whereas individual risk involves the possibility of casualty to an individual person. Such summaries of possible risks are important to the hazard assessors, urban planners and often to the general public, hence the need for standardized methods of presenting the results of risk analyses. Risk contours are often used; these are lines drawn on a map to connect points of equal risk (product of probability of an accident and the extent of damage caused or amount of casualties) as a result of any of the range of hazardous events which may occur. It is often useful to produce risk contours for various extents of damage as a result of different accidents. With LNG, there will be an overall risk contour map (which includes the possibility of domino effects); however, separate maps for envelopment by a flammable cloud, freezing, asphyxiation, and hazardous thermal radiation fluxes should also be produced. Another method of summary is the cumulative frequency or F-N curve. The F-N curve is a plot of accidents per year (frequency) versus number of deaths. As the number of deaths from an accident increases, the frequency of an accident with that severity decreases. On such curves plots are made for both natural and industrial accidents as well as the proposed facility in an effort to make a direct comparison of the extent of risk involved. A typical curve for an LNG facility is illustrated in Fig. 4. It is important to stress that a reasonable risk for a particular facility can be estimated. The estimation entails the following calculations. From failure rate data, typical flow rates, and probabilities of flows or volumes in tanks and lines, it is possible to assign probabilities for releases of different sizes. The consequences of release will then be largely determined by meteorological conditions, local terrain and probability of ignition. Probabilities for wind speed, wind direction and atmospheric stability can be obtained from

268

JSED

FOR

NUCLEAR

7\ IO DEATHS

100 EXPECTED

METEORS I

1000 FOR

10000 PARTICULAR

AREA

Fig. 4. F-N curve for an LNG facility.

those measurements made in previous years by the local weather offices. Using this data, it is possible to determine the distance and direction of movement of the vapour cloud. In travelling away from the spill site, several potential ignition sources may be encountered, in the form of plant equipment, persons nearby, or vehicles. A probability can be assigned to each known ignition source, which reflects its location (degree of concealment), the energy involved, and the percentage of time that it will act as an ignition source. There will then be two major sub-groups: cases where ignition takes place and cases where ignition does not take place, The consequences of either case have been discussed in earlier sections of this paper. There is a second condition for ignition and that is the mixture must be in the flammable range at the point where ignition is attempted. Having worked through all of the possible combinations and their associated probabilities, the last step is to generate the risk statistics by multiplying these probabilities by the damage which may be incurred. The damage, if measured in numbers of lives, will depend on the relative distance between the LNG facility and populated areas. If the prevailing winds are such that it would cause the vapour cloud to engulf a heavily populated area, then the risk is greatly increased. In estimating this risk, it is necessary to specify the time of day of the release as the probabilities for the number of people exposed to the hazard will differ during the course of a day. Average rates of deaths to both employees and general population can also be used as a basis of comparison. Fatal Accident Frequency Rates (FAFR) are often used for such comparisons.

269 CONCLUSIONS

The hazards posed by the production and storage of liquefied natural gas can be minimized to acceptable levels by good engineering design and a complete understanding of the physical properties of LNG and the response of materials of construction to these properties. Good engineering practice and the performance of hazard analyses in various forms has prevented a major accident involving LNG for over forty years. 1. The primary hazards posed by LNG are: high surface emissive power from a pool fire, extreme cold, asphyxiation, and explosion when the gas enters confined areas. 2. Releases of LNG do not combust via UVCE or BLEVE; however, pressurized flammable refrigerants on a liquefaction site may do so. 3. In modelling the dispersion of the heavy vapour the use of slab models and advanced slab models are recommended as they are simple, and provide acceptable results in comparison with the other uncertainties in the simulation. 4. In areas where RPT explosions can occur, they should be considered in the analysis as a mild blast hazard and a possible source of ignition. 5. In estimating the radiation fluxes from a pool fire, Method 3, involving a detailed configuration factor, provides the most accurate determinations. REFERENCES

7 8

9 10 11

Cox, R.A., Comer, P.J., Pyman, M.A.F. and Slater, D.H., 1979. Safety issues in the siting of modern LNG terminals, Gastech LNG/LPG Conf., Houston, p. 183. Wells, G.L., 1980. Safety in Process Plant Design. George Goodwin Ltd., Great Britain. Lawley, H.G., 1974. Operability studies and hazard analysis. Chem. Eng. Prog., Vol. 70 No. 4. and Hazan” Information Exchange Scheme, Institution Kletz, T.A., 1983. “Hazop of Chem. Engineers, Great Britain. Canadian Standards Association. Liquefied Natural Gas (LNG)-Production, Storage and Handling, Z276-M1981. Risk Analysis of Six Potentially Hazardous Industrial Objects in the Rijnmond Area - A Pilot Study. A Report to the Rijnmond Public Authority. D. Reidel Publishing Co., Dordrecht, Holland, 1982. Feldbauer, G.F., Heigel, J.J., McQueen, W., Whipp, R.H. and May, M.G., 1972. Spills of LNG on Water. Esso Research and Eng. Co. Report No EE 61D-72, Nov. McBride, W.C., 1981. A Denser than Air Vapor Dispersion Validation Study Using the ZEPHYR & DISCO Models. A Report by Energy Resources Co. to Exxon Research & Eng. Co. Taft, J.R., 1981. Simulations of Experimental Spills Using MARIAH Model. A Report by Deygon-Ra Inc. to Exxon Research & Eng. Co. SIGMET-N Model Description and Users’ Guide. SAI-236-81-232-LJ. Vols 1 & 2. Science Applications Inc., El Segundo, California, 1981. Schnatz, G. and Flothman, D., 1980. A K-model and its modification for the dispersion of heavy gases. In: S. Hartwig (Ed.), Heavy Gas and Risk Assessment. D. Reidel Publishing Co., Dordrecht, Holland.

12 13

14

15

16 17 18

19

20 21 22 23 24 25 26 27 28 29 30 31

32

33 34

LNG Terminal Risk Assessment Study for Oxnard, California. Science Applications Inc. (prepared for Western LNG Associates 1975). Cox, R.A. and Carpenter, R.J., 1980. Further development of a dense vapour cloud dispersion model for hazard analysis. In: S. Hartwig (Ed.), Heavy Gas and Risk Assessment. D. Reidel Publishing Co., Dordrecht, Holland. Fryer, L.S. and Kaiser, G.D., 1979. DENZ - A Computer Program for the Calculation of the Dispersion of Heavy Toxic or Explosive Gases in the Atmosphere, SRDR152 UKAEA. Germeles, A.E. and Drake, E.M., 1975. Gravity Spreading and Atmospheric Dispersion of LNG Vapour Clouds. 4th Int. Symp. on Transport of Hazardous Cargoes by Sea and Inland Waterways, Florida. Fay, J.A., 1980. Gravitational spread and dilution of heavy vapour clouds. 2nd Int. Symp. on Stratified Flows, Vol. 1. Tapir, Trondheim, Norway. Van Ulden, A.P., 1974. On the Spreading of a Heavy Gas Released Near the Ground. First Int. Loss Prevention Symp., The Hague, Netherlands. Picknett, R.G., 1978. Field Experiments on the Behaviour of Dense Gas Clouds. Salisbury. Wilts., Great Britain Chemical Defence Establishment, Porton Down, Report No. IL 1154/78, 1978. Colenbrander, G.W., 1980. A Mathematical Mode1 for the Transient Behaviour of Dense Vapour Clouds, 3rd Int. Symp. on Loss Prevention and Safety Promotion in the Process Industries, Basel, Switzerland. Stall, A.M. and Chianta, M.A., 1971. Heat transfer through fabrics as related to thermal injury, Trans. N.Y. Acad. Sci. Ser. 2, Vol. 33, No. 7, November. Lees, F.P., 1980. Loss Prevention in the Process Industries. Butterworth & Co. Ltd., London. Robertson, R.B., 1976. Spacing in chemical plant design against loss by fire. Process Industry Hazards, I. Chem. E. Symp. Series No. 47, p. 157. Roopchand, D.R., 1983. Prediction of the Effect From Spills of LNG. 33rd CSChe Conf., Toronto. Schneider, A.L., Lind, C.D. and Parnarouskis, M.C., 1979. US coast guard liquefied natural gas research at China Lake. Gastech LNG/LPG Conf. Houston, p. 215. Strehlow, R.A. and Baker, W.E., 1976. The characterization and evaluation of accidental explosions. Prog. Energy Combust. Sci. Vol. 2, No. 60, p. 27. Clancey, V.J., 1982. The Effects of Explosions, The Assessment of Major Hazards. I. Chem. E. Symp. Series 71. 1982 Manchester. Thomas, P.H., 1963. The Size of Flames from Natural Fires. 9th Int. Symp. on Combustion. Academic Press 1963. Mizner, G.A. and Eyre, J.A., 1982. Large-Scale LNG and LPG Pool Fires, the Assessment of Major Hazards. I. Chem. E. Symp. Series 71. 1982 Manchester. Welker, J.R. and Sliepcevich, C.M., 1966. Fire Technology, 2: 127. Raj, P.P.K., 1977. Calculation of Thermal Radiation Hazards from LNG Fires - A Review of the State-of-the-Art, AGA Transmission Conf. 1977 St. Louis, Missouri. Moorhouse, J. and Pritchard, M.J., 1982. Thermal radiation hazards from large scale pool fires and fireballs. The Assessment of Major Hazards, I. Chem. E. Symp. Series 7 1. Manchester. Napier, D.H. and Roopchand, D.R., 1984. Ignition Characterization of Rapid Phase Transition Explosions. Spring Technical Meeting of the Comb. Inst., Fredericton, Paper No. 17. Napier, D.H. and Roopchand, D.R., in press. Evaluation of ignition potential of rapid phase transition explosions in LNG/water systems. J. Inst. Energy. West, H.H., Pfenning, D.B. and Brown, L.E., 1979. A Systems Approach to LNG Fire Safety. Gastech LNG/LPG Conf., Houston p. 193.

- area measurement, see Fig. 2 area measurement, see Fig. 2 - area of spill : A’ - areas in Fig. 3, subscripts 1,2,3, 4 b - distance in configuration factor see Fig. 3 B - ratio see Fig. 3 factor see Fig. 3 c - distance in configuration C - ratio see Fig. 3 C’ - constant d - diameter of pool d’ - distance from edge of impounding area to target - inertia correction factor y - distance measurement see Fig. 3 g - acceleration due to gravity h - pool height H - bund height H, - minimum LNG pool thickness I - thermal radiation intensity 5 - radiation flux from a fire Eq. (11) - critical radiation intensity for pilot ign. 2 - critical radiation intensity for spontaneous ign. k” - thermal conductivity h, - constant k, - constant L - latent heat of vaporization - mass evaporated - mass rate of evaporation of LNG : .A#,- maximum mass rate of evaporation of LNG 4 - heat flux to pool Q - steady radiation flux eq. (8) Q’ - total heat radiated by flame r - pool radius r ’ - radius of pool at time t’ R - radial distance see Fig. 1 Rf - radiation intensity from flame R, - radiation intensity at distance x E - time, exposure, duration t’ - time of arrival of spreading pool at radius r’ t.5 - time duration after spill &I - time after spill at max. pool diameter Ta - initial ambient temperature of substrate Tb - boiling point of liquid Tf - flame temperature a

,

-

m2 z: m2 m m

m m m m/s2 m m F/m2 kW/m2 W/m2 W/m2 Worn K

J/kg kg kg/m2 s kg/m2 s W/m” kW/m2 kW m m

FWjrn’ kW/m’ S S S

“K

K K

272

- volume of liquid in pool - instantaneous spill volume continuous liquid spill rate V, X distance from fire centre line to target X1 - distance, see Fig. 2

m3 m3 m3/s m m

5 P 0 z;i -

m/s

V

Vi

ct’

A

thermal diffusivity of ground buoyancy factor emissivity of flame angle density L - liquid, W - water Stef~-~olt~man constant atmospheric t~ansmissivity configuration factor

kg/m3

kW/m2K4