Lightning phenomena and protection systems
3.5 Risk factors and protection
3.5.4.4 Lightning to tall structures
Only generic models can provide an estimation of the risk of lightning initiation from tall structures. The critical ambient thundercloud field necessary for upward leader inception decreases with the increase of structure height h because of the intensification of the local field. For structures with height h > 200 m, the critical ambient field becomes lower than 20 kV/m. The electric field Ecl created by a storm cloud at the earth’s surface before a lightning flash can approach this critical value (section 3.2.7). Because of space charge effects, the electric field increases with height above ground (section 3.2.7.1) and the field intensity before a lightning flash can be simultaneously Ecl = 5 kV/m at the earth’s surface and Ecr = 65 kV/m at 600 m. The electric field created by the storm cloud may be substantially less than the critical electric field needed for upward leader inception in the case of structures with
height h < 100 m, but for higher structures an upward leader may form and initiate a strike without a preceding downward negative leader. To incorporate this effect in the Petrov–Waters model for a downward flash requires the addition of the electric field Ecl created by the thunderstorm cloud to the electric field created by the downward negative leader (if present). Such a consideration allows an estimate to be made of the proportion of downward and upward flashes occurring to very high structures.
Application of the Petrov–Waters model then gives a maximum lateral displace-ment for a strike to the mast top that is approximated by the relationship:
max≈ 0.54[(h + 15)i0]2/3exp(αh) (3.70)
where, for example, for a storm cloud field of Ecl = 4.2 kV/m the mast height enhancement coefficient α= 0.0021. The coefficient α will of course increase with the intensity of the storm cloud field Ecl.
Since the risk factor of lightning strikes to a structure calculated from Equation (3.58) is:
R= π
∞
0
2max(i0, h)p(i0) di (3.71)
where p(i0)is the probability density function for the current amplitude distribution, then in the case when the lightning current amplitude is, for example, log-normally distributed this integral may be calculated analytically. Substituting the values for lateral displacement and calculating the integral we have for the collection area:
Ac = 153(h + 15)4/3exp(0.0042h) [m2, m] (3.72) Here, the values i0 = 31 kA and σ = 0.7368 for the mean and standard deviation of the peak current statistical distribution have been used [109]. The risk factor for a structure with height h can then be expressed as:
R= 153 × 10−6Ng(h+ 15)4/3exp(0.0042h) (3.73) In Table 3.8, the risk factor is presented for a ground flash density Ng= 1 km−2year−1. It is seen from the Table that the influence of the storm cloud field becomes important for structures with heights h > 100 m, indicating that the contribution of upward, structure-initiated flashes increases with the structure height.
The results describe well the observed data of Petrov and D’Alessandro [100, 112], including the probability of strikes to very high structures. It is known [66], that the average number of strikes per year to the 540 m television tower in Moscow
Table 3.8 Risk factor for a negative flash to a tall mast ( for Ng= 1)
h(m) 20 50 100 200 300 400 500
R(downward flashes) 0.02 0.04 0.09 0.2 0.3 0.5 0.6 R(all flashes) 0.02 0.05 0.13 0.45 1.2 2.5 5.1
is equal to R ≈ 30, using a value of ground flash density for the Moscow area of Ng∼= 3–4/km2/annum.
3.5.5 Protection of overhead power lines
The vulnerability of power systems to lightning is illustrated by considering that a direct lightning strike of median current (31 kA) to an 11 kV line (surge impedance 260 ) could produce a peak voltage of 4 MV. Modern lightning location techniques show that even in Northern Europe there were 540 000 ground strikes in 1992. On the UK 132 kV system, the lightning fault rate is between 0.3–1.5 faults/l00 km/year, with damage limited to 5–10 per cent of cases because of a high BIL and fault current interruption. 11 kV fault rates are lower, but the much larger system and lower BIL results in many more incidents and a higher consequential damage in 40 per cent of cases on lines and 100 per cent on cables and equipment.
Power supply plant such as pole-mounted transformers or substation equipment have to be protected against high voltage surges resulting from shielding failures and backflashovers. Lightning overvoltages can be predicted using electromagnetic transient programs for which the main parameters are the peak current, earth resis-tivity, line geometry, arrester characteristics and cable type. The return stroke current source is often modelled as a transmission line. Even SF6-pressurised GIS are vulner-able to very fast transients. The last two decades have seen advances in applications of metal-oxide surge arresters, solid-state overcurrent and distance protection, line sectionalisers and automatic reclose circuit breaker technology. ZnO gapless arresters in polymeric housings are effective provided that they are properly specified and are positioned close to the protected equipment. This will limit the overvoltage at the equipment terminals to Va+ 2ST , where Vais the residual voltage at the arrester, S is the surge steepness and T the transit time from arrester to the equipment terminals (see chapter 5, section (i)). The IEC Standard [113] gives important recommendations on the specification of arrester rated voltage, maximum continuous operating voltage, discharge current, energy rating and residual voltage.
The availability of online lightning tracking data offers transmission system engi-neers new possibilities for improved plant asset management and security of supply [21]. Early warning of severe storms is being improved, and the high precision of ground strike location within 500 m is often useful in fault location after a lightning event. Archival data of lightning activity also allows realistic estimates of risk factor and seasonal and geographical variations.
Where a direct strike is prevented successfully by shielding, overvoltages from backflashover to the phase conductor will be proportional to the peak current and are determined by the surge impedances of the tower and shield wire and the footing resistances [114]. The classical lattice diagram method remains useful to take account of factors causing variability in backflashover voltages, such as the steepness of the lightning current front, the point-on-wave of the supply voltage and the line shielding by the upward leader. A useful and substantial reduction of steepness and amplitude results from corona attenuation during the surge propagation along the line. Until recently, this effect has often been misinterpreted as either a reduction in
surge propagation velocity or a change in the surge impedance of the line. In fact, the corona discharge affects neither of these, but absorbs energy during the travelling voltage front because of the ionisation of the air around the conductor. Al-Tai et al.
[115] have shown how the corona attenuation factor can be predicted in terms of the line conductor geometry.
When transient analysis of the network provides an estimate of the overvoltage probability density p(V ), the level of protection required can be chosen on the basis of the risk factor R (section 3.5.3) and the risk of flashover RF(section 3.5.1). Close-in or direct strikes to substations must be prevented because the magnitude and rate of rise of the unattenuated overvoltage would disable protection by gaps or autoreclosure. Gary et al. [14] recommend a rolling sphere design of the shielding. Complete shielding will be achieved, even for the minimum observed lightning peak current of 2 kA, with a choice of sphere radius R= 15 m. Alternatively, a design current of 5 kA or greater would include 97 per cent of flashes, and is less onerous with its value R= 27 m.
For 0.2 flashes/year to the substation, this would be equivalent to one failure in 150 years.
The dependence of flashover voltages on the complex overvoltage shapes arising on power systems [116], and the consequent critical values of lightning peak current have been considered by Darveniza et al. [117] who deduced empirical equations for time to flashover for partly chopped impulses, and Hutzler and Gibert [118] who both modelled and tested backflashover impulse shapes. In more recent work, Haddad et al. [119] have quantified the flashover voltage for air gaps in parallel with ZnO surge arresters. Another factor is that following fault protection operation, an air gap can suffer significant arc erosion of the electrodes which will alter its volt–time breakdown curve.
3.5.6 Protection of electronic equipment 3.5.6.1 Strategy
Lightning electromagnetic pulses (LEMP) are a particular hazard for electronic systems. The indirect surges arising from inductive coupling and resistive voltage drops, and to a lesser degree capacitive and radiative coupling, are sufficient to cause severe damage; a pulse energy of less than 1 microjoule can easily destroy an inte-grated circuit. As in power equipment the requirements for the protection are speed and reliability in overvoltage control, survivability and system compatibility in nor-mal operation. The strategy of the protection is the same in both cases: where power engineers speak of main and backup protection, electronic engineers specify primary and secondary protection. A recent comprehensive review of the standards for the pro-tection of low voltage systems against lightning and other surges has been published by Hasse [120].
Three stages are generally necessary to achieve immunity to lightning problems:
(i) transient amplitudes must be controlled by means of the building design and equipment layout
(ii) the equipment must meet electromagnetic compatibility standards (iii) surge protector devices should be used to minimise let-through voltages.