Abstract. 1. Introduction. MEng Mechanical Engineering Schlumberger

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1 IBP3359_10 PREDICTION AND VERIFICATION OF TCP AND WIRELINE PERFORATING GUNSHOCK LOADS Carlos E. Baumann 1, Harvey A. R. Williams 2, Awad William 3, and Vincent Pequignot 4 Copyright 2010, Brazilian Petroleum, Gas and Biofuels Institute - IBP This technical paper was prepared for presentation at the Rio Oil & Gas Expo and Conference 2010, held on Sept 13-16, 2010, in Rio de Janeiro. This technical paper was selected for presentation by the Technical Committee of the event according to the information contained in the abstract submitted by the author(s). The contents of the technical paper, as presented, were not reviewed by IBP. The organizers are not supposed to translate or correct the submitted papers. The material as it is presented does not necessarily represent Brazilian Petroleum, Gas and Biofuels Institute s opinion, nor that of its Members or Representatives. Authors consent to the publication of this technical paper in the Rio Oil & Gas Expo and Conference 2010 Proceedings. Abstract When planning perforating operations, it is critical to minimize the risk of equipment damage caused by perforating gunshock loads, such as breakage of weak points when perforating with wireline conveyed guns and buckling of tubing and unset packers in TCP jobs. Perforating gunshock loads are produced by transient pressure waves in the wellbore fluid and stress waves in structural components. The magnitude and timing of pressure waves depends on design parameters that can be adjusted by the design engineer, such as the distance from the guns to packers and well bottom, loading of guns, number of shock absorbers used to prevent buckling and corkscrewed tubing, etc. This paper describes the use of a simulation program that predicts transient fluid pressure waves and the associated structural loads for most types of well perforating events. All relevant aspects of the well perforating event are modeled, including gun filling after firing, wellbore pressure waves and associated fluid movement, wellbore fluid re-pressurization with reservoir flow and through filling/debris subs, dynamic deformation of tubing and guns, and plastic deformation of shock absorbers. The predictive capabilities of this program are demonstrated by comparing fast gauge pressure data, shock absorbers deformation, and cable tension logs with the corresponding simulated values. Fast gauge pressure data shows that the predicted wellbore pressure transients are accurate both in magnitude and time. Residual deformation of crushable elements in shock absorbers indicate peak axial loads during the perforating job; such values are found to correlate well with predicated peak axial loads above the guns. Surface peak cable tension is also well predicted by this simulation tool. The main goal of the described simulation program is to minimize the risk of completion tools damage due to perforating gunshock loads. This program helps engineers to verify and/or optimize perforating jobs to reduce the risk of damaging costly equipment and time losses. 1. Introduction Perforating is the starting point for cased-hole completions. The goal is to create perforation tunnels with the necessary size, with minimum amounts of debris and crushed rock damage, and avoiding undesirable side effects such as excessive debris in the wellbore and damage to downhole equipment due to gunshock. Perforating gunshock can describe a broad range of phenomena, including but not limited to ballistic shock wave damage to downhole electronics and monitoring equipment, permanent deformation (corkscrewing) of tubing, and damage to packers; and specifically for wireline (WL) operations, gun jumping, breakage of the weakpoint and bird nesting of the conveyance cable. Through extensive field testing and laboratory experiments done in the last 15 years, driven by dynamic underbalance perforating for improved productivity, see Behrmann et al. (1997), Walton et al. (2001), and Baxter et al. (2009), we now understand the key mechanisms that control pressure waves in the wellbore during perforating events. Perforating pressure waves can create clean perforations with low skin; but under certain conditions, typically in high pressure wells, perforating pressure waves can produce equipment damage and other undesirable side effects such as sanding. Other works examining gunshock can be found in Schatz et al. (2004) and Canal et al. (2010); and gun jumping during underbalance perforating by Zazovsky et al. (2007). 1 Ph.D. Engineering Mechanics Schlumberger 2 Ph.D. Applied Mathematics Schlumberger 3 B.E. Electrical Engineering Schlumberger 4 MEng Mechanical Engineering Schlumberger

2 Herein we present details of a simulation model and case studies that demonstrate how key elements of gun shock are simulated. In subsection 1.1, we describe the background to the model; and in sections 2, 3, and 4 we present three case studies. 1.1 Background to the Model To understand perforating wellbore dynamics and gunshock, consider the detonation of shaped-explosive charges inside a hollow-carrier perforating gun. In about 100 microseconds, the shaped-explosive charge detonates forming the perforation tunnel. During this process, energy released from the explosive is converted into kinetic energy of the perforating jet, elasto-plastic deformation of the perforating gun hardware, and internal energy of the detonation gas. During the detonation of the charges, gun pressure is spatially non-uniform with a magnitude that can be larger than the wellbore pressure. When the perforating jets puncture the hollow carrier wall, the detonation gas inside the gun begins to interact with the wellbore fluid; this is the onset of wellbore fluid dynamics effects, when wellbore pressure transients begin. Wellbore fluid hydrodynamics is driven by interaction between three pressures sources: detonation gas pressure in the guns, wellbore fluid pressure around the guns, and formation pore pressure. In the context of gunshock analysis, dynamic underbalance refers to the fast drop in wellbore pressure caused by low gun pressure, and dynamic overbalance to a fast increase in wellbore pressure caused by high gun pressure. Under dynamic underbalance conditions, wellbore fluid flows into the guns, creating a large drop in wellbore fluid pressure because completion liquids have very low compressibility. The decompression of wellbore fluid around the guns drives two processes: pressure waves in the wellbore fluid propagating up and down the wellbore at the liquid's speed of sound; and fluid flow between the reservoir and the wellbore. The first of these effects produces gunshock forces, in essence a water-hammer effect. This paper describes a simulation tool capable of predicting transient fluid pressure waves and structural loads in the perforating system. Gunshock stresses or equivalent member loads can be large enough to damage the upper assembly and tubing. Compressive and tensile loading on the tubing can produce plastic deformation; a typical manifestation is permanent corkscrewed tubing joints due to helical buckling with plastic deformation of the tubing. The elasto-plastic simulation model computes the minimum axial compression load (critical load) needed to initiate plastic deformation on tubing buckled with casing support, taking into account all relevant changes in casing size, guns eccentricity, presence of centralizers, tubing pressure, guns weight, etc; the time dependent simulated maximum dynamic compressive load on the tubing is checked against the critical load. All relevant aspects of the well perforating event are modeled, including gun filling after firing, wellbore pressure waves and associated fluid movement, wellbore fluid re-pressurization with reservoir flow and from tubing flow/debris subs, elastic deformation of tubing and guns, plastic deformation of shock absorbers, etc. The model that computes the transient pressure field in the wellbore fluid and guns, also computes the transient gunstring deformation, velocity, acceleration, and stresses everywhere in all relevant structural components. A fully coupled fluid structure interaction strategy is used for this purpose. Prediction of pressure waves in wellbore fluids and stress loads on guns, tubing, and packers calls for a coupled fluid structure simulation. The program described in this paper simulates fluid flow along the wellbore annulus and in the guns with a finite element technique described by Baumann (1997), and Baumann and Oden (1999a, b) and (2000). This finite element technique is capable of delivering accurate and stable solutions to the Navier-Stokes equations. The structural simulation model is also based on finite elements. Both models are coupled through a high- order accurate time stepping discretization algorithm. The simulation model uses a special discretization method to evaluate reservoir flow accurately and efficiently, this is based on ideas borrowed from Baumann et al. (1999c). 2. TCP Job with 7.0-in Guns and 2-7/8-in Tubing The use of 3-1/2-in tubing or drill pipe is strongly recommended to perforate safely with 7.0-in guns. A gunstring with 2-7/8-in tubing joints can be used safely only when certain conditions are met; this requires a good combination of guns loading configuration and wellbore pressure. Here we analyze a perforation job done with 7.0-in guns, 12-spf 39-gram shaped charges 1, and 2-7/8-in tubing. Figure 1 shows the gunstring diagram and a bent tubing joint produced by the large compressive gunshock load generated during the perforating event. The net interval perforated was 90 ft, and the initial wellbore pressure at the firing head 5,250 psi. Figure 2 illustrates the transient pressure waves at 17, 34, 51, and 85 ms after the beginning of the detonation train. To simplify the explanation of this figure, wellbore pressure is plotted after removing the hydrostatic gradient, for this reason we see that wellbore pressure before firing is 5,250 psi throughout the length of the wellbore. 2

3 Figure 1: Gunstring diagram and bent 2-7/8-in tubing joint For the guns charge loading used, the gun pressure after firing is lower than the initial reservoir pressure. Therefore the guns become a pressure sink, this is observed in the wellbore pressure profile at 17 ms (red line), which shows a lower wellbore pressure around the guns as wellbore fluid started to fill the empty volume of the guns. Pressure waves in the wellbore annulus move at the speed of sound in the wellbore fluid (1,500 m/s in water), they reflect partially at every change in the fluid-occupied wellbore cross sectional area, and they may undergo a full reflection when reaching packers or the plug back. The wellbore pressure profile at 34 ms (green line) shows the lower pressure wave moving downwards with little change, whereas the pressure wave moving upwards (above the guns) has started to interact with the packer. Similarly, the pressure profile at 51 ms (blue line) shows the downwards moving pressure wave still moving with little change in amplitude, but the pressure wave moving upwards (on the upper side) has already started to interact with the packer, and the wellbore pressure at the packer is substantially lower than the initial minimum wellbore pressure of 4,125 psi above the guns. When pressure waves reach a blocking surface they double their amplitude (perfect reflection). This is clearly observed at 85 ms (purple line), the lower pressure wave has reflected on the well bottom, its original pressure amplitude (5,250 4,225) psi has now doubled to (5,250 3,200) psi. Similarly the pressure wave moving upwards has reflected on the packer and its amplitude has doubled from (5,250 4,125) psi to (5,250 3,000) psi. Figure 2: Wellbore pressure (without hydrostatic gradient) as a function of time 3

4 Figure 3 is a snapshot of the simulation program animation results at 308 ms; the plot on the left shows pressure along the wellbore (similar to Figure 2), and the plot on the right shows the time history of change in axial force right above the guns up to 308 ms, at the instant of peak dynamic compressive axial load. Figure 3: Wellbore pressure at 308 ms and tubing axial load history to 308 ms. The axial force on the tubing depends on the relative stiffness (elasticity) and mass of the guns and the tubing between the guns and the packer. The time dependent axial force on the tubing is not the same as the net fluid/gas force acting on the guns because of the large mass / inertia of the guns, which absorb or release energy during the dynamic event. This is illustrated in Figure 4, showing the time history of change in resultant axial force on the guns (blue), and the change in dynamic force on the tubing right above the guns (red). In Figure 4, we see that during the first 100ms the net applied force on the guns is much larger than the axial force on the tubing, this is because prior to detonation the guns are stationary, and even when a large load is applied on them, only a fraction of this load is transmitted to the tubing because the guns absorb the load impulse while accelerating from rest. Conversely, at later times in the dynamic response (after 120 ms in Figure 4), the dynamic forces on the tubing can be much larger than the applied load on the guns, even when the net force on the guns is very low, as the guns continue oscillating they can exert large forces on the tubing if their combined momentum is sufficiently large. The latter effect is clearly shown in Figure 4 after 300 ms, where we see that the net applied force on the guns is very low, but the peaks of tubing axial force continue to be large. Figure 4: Above guns incremental tubing force history to 720 ms. 4

5 Figure 4 also shows shaded areas indicating periods where helical buckling with plastic deformation occurs. Tubing can buckle inside the casing without any permanent deformation, as long as the peak stress in the tubing remains below the yield stress. But even with the lateral support of the casing, if the compressive axial load is larger enough the peak stress in the buckled tubing can exceed the yield stress, and this will be visible as permanent corkscrewed tubing after pulling out of hole (POOH), as illustrated in Figure 1. Elastic sinusoidal and helical buckling of tubing inside casings has been described extensively in the literature; see for example the seminal work of Lubinski et al. (1962). Extended reviews of tubing buckling were published by Cunha (2003), and Aaseen and Aadnoy (2003). The post-buckling effect of friction between tubing and casing was studied by Gao and Miska (2009). The axial load that produces elastic buckling (whether sinusoidal or helical supported by the casing) is typically much lower than the axial load needed to produce permanent deformation of the tubing. Figure 4 shows highlights the time interval where the compression load on the tubing produces plastic deformation; 30k lbf is the axial load that produces the first traces of plastic deformation on the surface of the tubing taking into account tubing yield stress, tubing and casing size and total guns buoyed weight, and tubing inner and outer pressure. In this case at 30k lbf the residual deformation of the tubing would be negligible. As the tubing is subject to larger axial loads, more plastic deformation occurs, and more residual deformation (corkscrewed tubing) will be noticeable after POOH. Figure 5 shows the change of force on the packer in time up to 720 ms. The applied force curve (blue) is an ideal or reference force obtained by adding the net force applied on the guns (fluid pressure force) and the net change in force over the packer due to the change in wellbore pressure below the packer. The dynamic force (red) curve represents the actual predicted change in packer force due to the perforating event, this force is the sum of the dynamic force on the tubing right below the packer, and the net change in force over the packer due to the change in wellbore pressure below the packer, the second component of the force is identical to the second component of the applied force curve (blue). As explained before, the dynamic force on the tubing right below the packer can be very different from the net applied force on the guns; this is why the blue and red curves differ. Note that the dynamic force on the packer remains zero during the first ~ 25 ms; this is the time that takes the stress wave to travel from the guns to the packer. At later times the changes in the dynamic force on the packer are much larger than the changes in the applied force, this is because the guns continue moving / vibrating even when the time dependent change in excitation force is relatively small. The peak change in dynamic total force on packers and plugs has to be checked to prevent the occurrence of unset packers. Figure 5: Packer incremental dynamic force history up to 720 ms. 3. Wireline Perforating Job with 4-1/2-in Guns This section analyzes a wireline perforation job executed without any delays or incidents. High-speed pressure gauge and surface cable tension data are available for comparison with simulation results. Net interval perforated 84 ft, 4-1/2-in guns, 12-spf 22-gram shaped charges 2. Initial guns weight in water 1750 lbs. Initial wellbore pressure at the firing head 9,200 psi. String: cable 7-48 ZA XXS (15,000 lbf), weak point (rating 6,850 8,500 lbf), WPSA, CAL-B, four 20-ft 4-1/2-in guns, and 2-7/8-in fast-gauge carrier. 5

6 The structural dynamics code takes into account all relevant changes in structural stiffness and mass, cable movement / deformation / tension are simulated all the way to surface. Also elements such as shock absorbers are included, Figure 6 (left). Figure 6 (right) shows the fast-gauge pressure data (red) and the simulated gauge pressure (green). The pressure amplitude is simulated accurately; there is only a small time shift of ~ 0.05 sec that could be attributed to uncertainties in the location of a casing size change. Figure 6: WPSA crushable element (left); high-speed gauge pressure and simulated pressure (right). Figure 7 (left) shows the relative magnitude of the perforating incremental force acting on the guns (blue), the change in cable tension at the weak point (red), and the cable tension at surface (green) that remains constant during the time period shown on this figure. Note the relative magnitude of the forces, the large initial lifting force on the guns is followed by a large downward pull that lasts longer, and eventually by a relatively small amplitude force that oscillates around zero (relative to the pre-firing state). The net change in cable tension is much smaller, this can be more clearly observed on Figure 7 (right). Figure 7 (right) shows peak cable tension at the weak point ~ 4,000 lbf; this value can be correlated with the maximum deformation of the crushable element shown in Figure 6 (left). The cable tension reduces after approximately 0.2s, as wellbore pressure recovers rapidly due to the combined effect of wellbore storage and reservoir flow. Figure 7: Predicted cable tension and load on guns (left); measured and simulated cable tension at surface (right) On Figure 7 (right) we also see the change in cable tension at the surface. The 2.1 s delay is due to the finite speed of sound in the cable. This is the time taken by any load perturbation at the cable bottom end to reach the surface. 6

7 WFDD data provides min and max cable tension at the surface at a 1 Hz frequency. Figure 7 (right) shows the min and max surface tension with purple and blue markers, respectively. The time shift has been set to make the simulated surface tension compatible with the WFDD data. Surface tension before shooting 7,500 lbf, and the max surface tension recorded while shooting 11,300 lbf, this value is only slightly higher than the predicted peak tension at surface. 4. Deepwater TCP Job with 6-5/8-in Guns and 3-1/2-in Tubing Here we analyze a perforation job done with 6-5/8-in hollow carrier guns, 18-spf 34-gram shaped charges 3, and 3-1/2-in tubing. This case is of interest because permanent deformation in a few 3-1/2-in tubing joints was very light, after POOH the tubing joints were rolled and it was observed that a few were bent approximately 1 to 2 inches, indicating incipient plastic deformation on the tubing s surface. Figure 8 shows the gunstring diagram, the explosively initiated vertical shock absorber (SXVA), and one of the two crushable elements. The net interval perforated was 48 ft with a distance from top shot to packer of 460 ft. Initial wellbore pressure at the firing head was 17,970 psi with a 14.5 ppg completion fluid. Figure 8: Gunstring diagram (left), SXVA (center), and SXVA s deformed crushable element (right). Figure 9 shows the high-speed pressure signal recorded by the high-speed gauge in the firing head (red), and the simulated wellbore pressure at the same location. We see good agreement for: the initial dynamic underbalance amplitude, the timing of the wave reflected at the packer and its amplitude when it reaches the firing head, and also the timing and amplitude of the second reflection. These are important simulation quality measures because their magnitude and timing are directly related to the magnitude of gunshock loads. Figure 9: High-speed pressure and simulated pressure at the firing head. Figure 10. Bull nose movement; positive indicates displacement downhole. 7

8 Figure 9 shows a good asymptotic convergence of measured and simulated pressure after 0.75 s, this is a good indicator that the reservoir effect (surge) is well modeled. A good match of reservoir flow / surge is important because in typical TCP jobs a strong reservoir surge decreases the magnitude of dynamic underbalance and the associated gunshock loading. Figure 10 shows the movement of the bull nose and the maximum predicted gunstring stretch that is 11 inches when measured from the initial position at rest, before firing. Comparisons of gunstring acceleration and acceleration histories obtained with IES gauges (Low G) indicate that both the magnitude and frequency of the gunstring movement are accurately predicted. Compression and tension loads on the tubing can produce plastic deformation; a typical manifestation is permanent corkscrewed tubing joints due to helical buckling with plastic deformation of the tubing. The elasto-plastic simulation model used here computes the minimum axial compression load (critical load) needed to initiate plastic deformation on tubing buckled with casing support, taking into account all relevant changes in casing size, guns eccentricity, presence of centralizers, tubing pressure, guns weight, etc. Figure 11 is a snapshot of the simulation program animation results at 435 ms; the plot on the left shows pressure along the wellbore at 435 ms, and the plot on the right shows the time history of change in axial force right above the guns up to 435 ms, when the peak dynamic compressive axial load occurs. Figure 11 also shows a shaded area indicating levels of compression axial loading that produces helical buckling with plastic deformation of the tubing. Figure 11: Wellbore pressure at 435 ms (left) and tubing axial load history up to 435 ms (right). The time dependent axial force on the tubing is typically different from the net fluid/gas force acting on the guns because of the large mass / inertia of the guns, which absorb or release energy during the dynamic event. This is illustrated in Figure 12 (left) showing the time history of change in resultant axial force on the guns (red), and the change in dynamic axial force on the tubing right above the shock absorbers (blue). In Figure 12 (left), we see that immediately after detonation (up to 30 ms) the net applied force on the guns is much larger than the axial force on the tubing; this is because at the beginning the guns are stationary, and even when a large load is applied on them, only a fraction of this load is transmitted to the tubing because the guns mass / inertia absorbs the load while accelerating from rest. Conversely, at later times (after 300 ms in Figure 12) the dynamic forces can be much larger than the applied load on the guns, and even when the net force on the guns is very low, as the guns continue moving / vibrating they can exert large forces on the upper assembly if their combined momentum is sufficiently large. The latter effect is clearly shown in Figure 12 after 300 ms, where we see that the net applied force on the guns is very low, but the peaks of tubing axial force continue to be large, in fact, the peak compressive load on the tubing (70k lbf ) occurs at 430 ms. In Figure 12 (left) we see that the peak compressive load on the tubing occurs at 430ms; the shaded areas indicate periods where helical buckling with plastic deformation occurs. The peak compression load predicted at the 8

9 explosively actuated shock absorbers is approximately 70k lbf; with this peak we can determine the residual deformation of the crushable elements using the force vs. deformation chart shown on Figure 12 (right). From this chart we extract a residual deformation of approximately 2 inches; this value is very close to the SXVA crushable elements deformation measured after POOH. Figure 12: Incremental tubing force above guns (left). SXVA crushable element force vs. deformation (right) 5. Conclusions This paper describes the latest advances in simulation software used for predicting gunshock loads on completion equipment. The simulation program described predicts transient fluid pressure waves and the associated structural loads for most types of well perforating events. All relevant aspects of well perforating are modeled, including gun filling after firing, wellbore pressure waves and associated fluid movement, wellbore fluid repressurization with reservoir flow and fill subs, dynamic deformation of tubing and guns, and plastic deformation of shock absorbers. The software allows the analysis of perforation job designs and associated gunshock loads. The effect of all perforation design elements and dimensions can be investigated, for example: Position of packers Size and length of conveyance Length of perforated interval and reservoir properties Guns and shaped-charges sizes Guns loading strategy Wellbore fluid properties Type and number of shock absorbers The predictive capabilities of this program were demonstrated by comparing fast gauge pressure data, shock absorbers deformation, and cable tension logs with the corresponding simulated values. The end goal of the simulation program described is to minimize the risk of completion tools damage due to perforating gunshock loads, thus reducing the risk of costly equipment damage and time losses. 6. Acknowledgements The authors would like to thank Schlumberger for setting high technical standards and for supporting this work. In particular, we would like to thank Andy Martin, Alan Salsman, Brenden Grove, Larry Behrmann, and Riley Suffridge for reviewing this work; and also Alex Moody-Stuart and Ram Shenoy for their support of Perforating Research activities. We also want to thank Edison Bustillos Pesantes for his valuable contribution specifying actual perforating job conditions, and Tom Stevenson for developing and maintaining the software s user interface. 9

10 7. References AASEN, J.A., AADNOY, B.S. Buckling models revisited. SPE 77245, BAUMANN, C.E. A new hp-adaptive Discontinuous Finite Element Method for Computational Fluid Dynamics, Engineering Mechanics Dissertation, The University of Texas at Austin, August BAUMANN, C.E., ODEN, J. T. A discontinuous hp finite element method for the solution of the Euler and Navier- Stokes equations. Int. Journal for Numerical Methods in Fluids, Volume: 31, Issue: 1, Pages: 79-95, 1999a. BAUMANN, C.E., ODEN, J. T. A new discontinuous hp Galerkin method for convection-diffusion problems. Computer Methods in Applied Mechanics and Engineering 175, pages , 1999b. BAUMANN, C.E., PRICE, H.S., REDDY, M.P., and THUREN, J.B. Full field pressure simulations using a very accurate, yet inexpensive well model. SPE MS, SPE Annual Technical Conference and Exhibition, 1999c. BAUMANN, C.E., ODEN, J. T. An adaptive-order discontinuous Galerkin method for the solution of the Euler equations of gas dynamics. Int. Journal for Num. Methods in Eng., Volume: 47, Issue: 1-3, Pages: 61-73, BAXTER, D., BERHMANN, L.A., GROVE, B., WILLIAMS, H., HEILAND, J., HONG, L.J., KHONG, C.K., MARTIN, A., MISHRA, V.K., MUNRO, J., PIZZOLANTE, I., SAFIIN, N., and SUPPIAH, R.R. Perforating when failure is the objective. Oilfield Review 21, #3, BERHMANN, L.A., Li, J.L., Li, H., Borehole dynamics during underbalance perforating. SPE 38139, CANAL, A.C., MILETTO, P., SCHOENER-SCOTT, M.F., MEDEIROS, J., BARLOW, D. Predicting pressure behavior and dynamic shock loads on completion hardware during perforating. OTC 21059, CUNHA, J.C. Buckling of tubulars inside wellbores. A review of recent theoretical and experimental works. SPE 80944, GAO, G., MISKA, S. Effects of friction on post-buckling behavior and axial load transfer in a horizontal well. SPE , LUBINSKI, A., ALTHOUSE, W.S., LOGAN, J.L. Helical buckling of tubing sealed in packers. Petroleum Transactions, p , ODEN, J. T., BAUMANN, C.E. A Conservative Discontinuous Galerkin Method for Convection-Diffusion and Navier Stokes Problems. Discontinuous Galerkin Methods - Theory, Computation and Applications, Lecture Notes in Computational Science and Engineering, Vol. 11, Springer Verlag, SCHATZ, J.F., FOLSE, K.C., DUPORNT, R. High-speed pressure and accelerometer measurements characterize dynamic behavior during perforating events in deepwater Gulf of Mexico. SPE 90042, SPE Annual Technical Conference and Exhibition, WALTON, I.C., JOHNSON, A.B., BEHRMANN, L.A., ATWOOD, D.C. Laboratory experiments provide new insights into underbalance perforating. SPE 71642, SPE Annual Technical Conference and Exhibition, ZAZOVSKY, A., Zhan, L., KUSUMADJAJA, A., CHEWYAM, J.K. A gun jump model for underbalance perforating. SPE , SPE Europe/EAGE Annual conference and Exhibition, Deep penetrator marketed under the name PowerJet Omega Deep penetrator marketed under the name PowerJet Omega Big-hole charge marketed under the name PowerFlow