It is not known whether there is a threshold to the dose required to produce genetic effects, since the data is very scant. It is difficult to obtain

53 DNA / >4 в—sr (©) r—————- —’O’ B—SC "^P B—SZ_ ^P в—sr P An incoming photon (y) collides with an orbital electron (-) of one of the atoms of tissue. A DNA molecule (BSP BSP. . J rests nearby. (The letters stand for subunits of the large DNA mole­cule: В = a base, S = a sugar, deoxyribose, and P — a phosphate J

An incoming electron (-) (either a primary beta ray or the secondary product of an X or gamma ray) ricochets among tissue atoms, ionizing some of them.

results at very low dose rates, although present evidence would seem to infer that some threshold exists (3a, b).

In the meantime a linear relationship must be assumed between damage and dose. This is pessimistic and is likely to overestimate damage, although popular science writers would take the opposite view.

Radiation affects genes and gives rise to mutations, a large number of which may be undesirable (99% has been suggested). A normal person naturally experiences mutation of genes due to natural background radio­active and chemical disturbances, at a rate of about 2% to an individual. Thus for each child there is a natural 1 in 50 chance that it carries a new mutation. Extra radiation would increase this chance slowly.

Figure 5.2 shows the range of damage effects as a function of dosage. For a whole body exposure of 2 rem, the normal rate of mutation is increased by between 1.5 and 6.5% of the original value. For a whole body dose of 5 rem the normal rate is increased by between 7 and 17%, and not until a whole body exposure of over 30 rem is experienced, is the normal rate of mutation is likely to double to a total of 4% per individual (4).

High pressures of the energy release are assumed to give rise to shock damage to the vessel in the radial and bottom head directions. The only experimental information comes from modeling and the use of scaled chem­ical explosive tests using TNT and pentolite. The models used included some welded, nozzled, and noncylindrical vessels to give some realism to the yield calculations (22). Some additional information may be gained from the SL-1 accident and the consequent residual strain which was measured in the vessel and core debris (25). The information is relevant, despite the fact that the SL-1 was a thermal light-water-cooled system.

Figure 5.8 shows a comparison to be expected from nuclear and chemical explosive results. The chemical explosion is faster and has a higher initial shock but a lower residual pressure than the nuclear explosion with the same integrated energy value.

Thus it is pessimistic to use the chemical explosion correlations produced for the initial radial shock deformation of the vessel. The residual pressures resulting from the explosion should be handled differently, and Section

5.5.2.2 indicates a method of reverting to a use of the total energy release for a hydrodynamic description of the excursion.

The nuclear power plant is seen as an answer to the air pollution problem, since in operation its only emissions would comprise very low radioactive waste gases with a radioactive level comparable to natural background activity. The nuclear power plant has none of the chemical emissions which may be released in large quantities from a fossil-fueled plant.

The matter of radioactive emission from an operating power plant is dealt with in Chapter 5. It should also be noted that modern nuclear plants are now being designed for essentially zero release.

Although this book deals with reactor safety and places emphasis on licensing only as a means of showing that safety is established, it is never­theless appropriate here to comment outside these terms of reference. By detailing possible emissions from fossil-fuel plants, it is possible to view the very minor nuclear plant emissions in a more balanced light during the licensing process.

The boundaries between different regions of the boiling channel are specified by experimental correlations. They are required in modeling be­cause we need to know where the heat-transfer mechanism changes from convective to bulk-boiling and where voiding occurs.

In a steam generator problem, it is necessary to integrate the heat transfer correctly and to obtain accurate time constants. The latter depend entirely on the heat-transfer coefficients (see Section 1.3.1.2). In a water-cooled core channel it is necessary to know the total core voidage, because its reactivity effect may be large (due to a removal of moderator) and fast (due to flash­ing or collapse of vapor voids).

It is possible, knowing the position of the boundary between the highly and slightly subcooled regions, to simplify the channel representation in a PWR. Here it is assumed that convective heat transfer exists below the boundary and boiling transfer exists above it. The equations for the mixed enthalpy of the fluid are used, and another experimental correlation is used to predict total subcooled voidage in the channel. The analysis for a channel that includes subcooled nucleate boiling and the experimental correlations used to predict boiling boundaries and voidage have been given by Bow­ring (9) for water systems operating in certain temperature ranges. If how­ever a channel includes bulk boiling, the small subcooled boiling range may be insignificant and the only boundary of interest is that marked by the saturation temperature or a certain superheat above this saturation temperature.

Work on this system has now largely ceased in the United States and in Europe. However, it does have very different characteristics from either of the other systems and is worth including (Fig. 2.3).

The design considered could employ either stainless steel clad pins or stainless steel cermet. The coolant is supercritical steam; therefore the

system pressure is very high in the range of 3700 psia. Temperatures are very close to those of the LMFBR system.

With steam coolant a direct cycle system is the natural design configura­tion, although steam cycle efficiencies may demand one or more stages of reheat. Table 2.1 shows the main characteristics of the supercritical steam — cooled system.

A reduced flow rate results in reduced heat removal and eventually damage to the fuel pin. If the pin is sodium-bonded then this damage will first appear as bond vaporization with consequent insulation of the fuel. Thus failure of the fuel pin might be considered to start when the fuel surface heat flux exceeds a critical value for sodium vaporization. This is rather similar to pool boiling and the critical heat flux (CHF) is given by Eq. (3.1).

CHF = 534,000 р0Л1 Btu/hr-ft2 (3.1)

where p is pressure (psia) (lb).

Such a failure criterion corresponds to PWR burn-out conditions. The criterion can be modified by local conditions in any particular design and may possibly be correlated to fuel surface temperature for a particular bond pressure. In one design in which it was calculated that when the fuel surface temperature exceeded 1900°F, the critical heat flux was attained, this sur­face temperature could itself be used as a failure criterion.

Bond boiling would effectively insulate the hot fuel until fuel melting had occurred. The molten fuel then might come into contact with the cladding, so that although vaporization of the bond might not burst the cladding, molten fuel in contact with the cladding or fuel vaporization would.

This failure criterion does not however apply to gas-bonded oxide fuel pins and other failure criteria are required.

The other possibility is that the bubble in the channel would be relatively large and would either lodge somewhere or would pass through the core transiently insulating the fuel pin as it passes through.

Studies have shown that the passage of single bubbles at normal sodium velocities, assuming complete insulation of the pin surface under the bubble, results in perturbations of tens of degrees rather than hundreds. Thus no sudden failure is likely (5b). If the bubble were to lodge in a single position for a comparatively long time, then the surface temperature could increase, unless the axial conduction or circumferential conduction paths were ade­quate in removing the heat from the surface under the bubble.

Axial conduction is not efficient and thus most heat is removed circum­ferentially through the cladding from beneath the bubble. Whether this latter path is efficient or not depends on the circumferential coverage of the bubble. For small bubbles the heat is removed through the cladding faster than it is input from the fuel within the cladding, but for larger bubbles the. heat input is the faster process. Thus there is a critical coverage which may be obtained from a comparison of the time constants for heat input into the cladding and heat transport circumferentially through the cladding.

For a bubble which covers an angle в of the fuel pin surface and assuming no heat removal (see Fig. 4.3), the time constant for the input of heat into the cladding below the bubble is given by

Tin = fefCf&dCcd^2 fa)l(p{C{R + 2gcdccdt) K{(RlV 3 + b) (4.9)

Assuming no heat input, the time constant for heat removal along the cladd­ing in a circumferential direction is given by

Tout = (QcicciR*e*)/4Kci (4.10)

The critical angle of coverage вст is defined as that angle when these time constants are equal. Below this angle of coverage the cladding surface temperature rises only moderately above normal values, whereas above this value cladding temperatures rise considerably more. The critical angle 6CI is given by

всг = (4 e(CfKcd dt)1/2/[(fttc(R + 2pcdccdl) Kt(R/f3 + P)]1/2 (4.11)

An approximate expression may be derived as

0cr = 0.268(8A’cjt/A’fT?)172

For a fuel pin of 0.25-in. diameter, cladding thickness of 0.0125 in., and with oxide fuel clad in stainless steel, the critical angle is 94°.

The temperature rises beneath such a bubble may be calculated using one of the three dimensional heat transfer codes which are available, TOSS or TRUMP (see the Appendix). The models used in these codes are the basic heat transfer equations detailed in Section 1.2 but arranged on a mesh that connects in three dimensions, making computer solution a ne­cessity. Many of the difficulties involved in these calculations are connected with the spatial finite difference approximations used to represent conduc­tion from one point in the mesh to another.

If the complete pin were blanketed by a bubble and the angle of coverage were 360°, then axial conduction would become important. Studies have shown that, for the above fuel pin, approximately an inch would have to be insulated by the bubble before failure of the fuel pin cladding could occur. For bubbles less than 1 in. in extension only moderate cladding temperature rises would result.

a. Coolant contamination. Coolant contamination persisted. It was eventually cleared by installing an external cold trap of much larger capacity, by adding plugging meters to detect impurity, and by additional pipework for balancing of cover gas volumes. Some of the now unused purity lines still remain blocked.

b. Design changes. Because of design changes, the original zero power physics calculations made on the ZEUS pile were invalidated. Therefore the reactor had to operate at 1 MW until it provided its own physics data to provide confidence for its high power operation.

c. Reduced stability. At high burn-up at 45 MW operation, less stability margin was apparent. This was apparently due to fuel expansion within cracks at the higher burn-ups, thus reducing the axial expansion feedbacks that were a negative contribution. This was no problem, but it pointed to the need for reactivity feedback measurements with adequate perturbations. The DFR oscillator had а І0 perturbation that was far too small for this investigation of feedback reactivity.

d. Radial blanket overheating. Due to maldistributed flow in the radial breeder at low flows, the blanket was severely damaged by stationary bubbles caught in buoyant down-flow conditions. Six hundred elements were removed, 200 only after cutting swollen sections. In fact, insufficient attention had been paid to the hydraulic design of the blanket sections, as these sections had low power ratings.

e. Core fuel pin damage. As entrained gas bubbles during start-up may adhere to unwetted fuel pins and then, due to this adherence enhanced by their buoyancy, stay long enough to blanket the pin as the power is raised, some fuel pins have been subjected to failure. This failure mechanism is troublesome in reactor operation but has been valuable in the sense that in no cases has there been any evidence that pin failure has spread or propa­gated (see Section 4.4.2) (26).

f. Primary circuit leak. A small leak in the primary circuit caused a one — year operating delay in order to find and repair the leak. The location of the leak was very difficult to determine, because the reactor is a 24 loop system any one of which could have caused leaking sodium to appear at the point where it was detected in the leak jacket.

First the cover gas was pressurized without affecting the leak. Then 10 gm of gold tracer were introduced into the primary and the tracer’s presence in the leak jacket subsequently confirmed the leak presence. Successive isolation of the leak location was made possible by adding helium to the cover gas and lowering the sodium level slowly to try to determine when the leakage stopped. This confirmed that it was in one of 24 bottom outlet pipes. Then, with a pressurized leak jacket and manual movement of the heat exchangers attached to each outlet pipe in turn, inward bubbling from the leak jacket was detected.

The failure was finally determined to be a fatigue failure of a bad and misaligned weld combined with thermal stressing due to cold NaK entering the hot outlet stream from a subsidiary purity line just upstream of the weld. Neither weld nor thermal stressing was sufficient to cause the failure alone. The purity line was removed and the weld remade to solve the problem.

Whether a wind condition can disperse a radioactive plume depends on whether conditions will allow the plume to be swept away, to spread out, and to eventually descend. These conditions are tied to the atmospheric stability and the temperature gradient in the atmosphere (11). A number of different thermal conditions contribute to this atmospheric gradient:

(a) Potential temperature decrease (lapse rate) with height is due to a temperature decrease as the pressure decreases. Assuming a dry adiabatic lapse rate without heat exchange to other air masses, there would be an upward cooling rate of 5.4°F/1000 ft. Such an adiabatic temperature con­dition ensures that the air will be unstable, because daytime hotter air nearer the earth is constantly changing position with upper colder air.

(b) The normal environmental lapse rate is not adiabatic, and it has some heat exchange, resulting in a lapse rate of 3.5°F/1000 ft. Superadiabatic rates give greater cooling with altitude than the normal environmental lapse rate. Also possible are subadiabatic rates giving a lesser cooling gra­dient. Both conditions are unstable.

(c) In some conditions the gradient can reverse itself, giving rise to inver­sion conditions, where the temperature increases with height. In inversion conditions the air layers are stable, cooler air remains below, and warmer air remains above. Such morning conditions can remain until the sun warms the lower regions to reverse the gradient and make the air mass unstable.

(d) In practice there may be several layers, each with a different thermal gradient.

Such wind speed and thermal stability considerations determine the wind — induced dispersal of any radioactive cloud following an inadvertent release of effluent. The wind conditions have to be characterized in terms of stability, persistence, shear, and direction. This characterization has been performed by Pasquill (12) and is used in dispersion calculations.

5.6.2.1 Blast Shielding

To allow the vessel to deform radially but not to such an extent that it damages the reactor cavity walls or thereby itself, blast shielding may be used effectively to strengthen the radial wall.

Figure 5.17 illustrates the radial cross section of shielding of a typical LMFBR. The shielding is provided for radiation reduction, and it may therefore be composed of serpentine concrete encased in steel. However,

the total radial thickness of mild steel (7.5 in. in this case) forms a very efficient and effective blast shield, even taking no credit for other materials in the shielding system.

The analysis of the blast shield is performed using the Cole charge weight correlation (24):

/ = 1.46 W°-33(W°-33IR)0-9 (5.29)

where / is the specific impulse (lb-sec/in.2), W is the charge weight (lb TNT), and R is the reactor shield radius (ft).

The kinetic energy is

KE = (IAfjlM (5.30)

and the energy of strain is

SE = In do&t (5.31)

where A is the inner surface of the unit annulus (in.2), M is the annulus mass (slugs), ё is the radial elongation (in.), t is the annulus thickness (in.), and <ra is the average stress during the strain (lb/in.2).

For the blast shield to perform as required the strain energy must be greater than or equal to the kinetic energy. Assuming an elongation before rupture of a given amount and summing the strain energies for each annulus of the blast shield, this sum can be compared to the calculated kinetic energy to determine whether the blast shielding is adequate.

To satisfy the model, the nitrogen-cooling annuli and the concrete annuli are assumed to have been collapsed without significant energy absorption. The shield of Fig. 5.17 can accommodate 30001b of TNT assuming an elongation SR/R of 0.3. Thus it can be seen that the inclusion of radial radiation shielding will usually provide more than adequate blast shielding at the same time.