No. Acceleration is a function of pressure over time. Longer barrels equal more 'T'.Originally Posted by Seth Harness
|
||||||||||||||||||||||||||||||||||||||
Senior Member
No. Acceleration is a function of pressure over time. Longer barrels equal more 'T'.
Senior Member
???
I'm not seeing how acceleration has anything to do with this. Pressures will be the same given the same chamber dimensions, powder charges and powder burn rates.
Pressures will be identical until the bullet passes the port. Then the shorter length system will bleed off pressure sooner, but at the point it's bleeding it, the pressure is significantly higher compared to the port location on a rifle length system where the pressure bleeds off. In a rifle length system the volume that the gas occupies is larger than in the shorter, carbine length system. Larger volumes result in lower port pressures.
Member
That makes a whole heap of sense.
So, is there a way to mitigate the higher pressure? Not increasing lock time and slowing the cycle rate but decrease the pressure.
Senior Member
The port will see the high pressure unless it's bled off earlier. In so doing, you'd need to bleed it off and vent it out to the atmosphere, you wouldn't want to direct it back to the BCG. But in so doing, I would expect velocity to take a hit.
The amount of gas that gets directed back to the BCG is related to the port diameter. Make it too small and you won't send enough gas back to cycle the action. Make it too big and the system is overgassed. Run your weapon hard and that gas port, which was properly sized when new, will erode and enlarge. Over time it will send more and more gas back to the BCG.
New weapons cycle around 650-750 rpm IIRC. As the port erodes cyclic rate can jump to over 1,000 rpm. That's where things start to get dicey WRT reliability.
I wonder if someone has ever concocted a bleed valve arrangment to regulate port pressure? if so, where did they vent the gas, forward maybe? I also wonder why the M4 designers haven't designed an adjustable gas valve to regulate the amount of gas directed back to th BCG. I'm pretty sure JP and MGI both make adjustable gas systems, but I'm not sure how effective and field-adjustable friendly they are. I think if someone could make one that could be easily adjusted in the field, that would be a huge boon to the M4 system.
Member
We(MGI) have the regulator on the gas tube currently, but we have a design for a newer product, which is more comprehensive in nature, but we just haven't been able to get it out the door, with all the other work we're doing with the MGI Modular Hydra AR15. Hopefully, we'll get it out before summer.
Getting back to the bolt breaking issue, of course the gas system is going to have an effect on this in various ways. I'm so glad that finally people are observing that just because a carbine "comes that way from the Colt factory", doesn't mean that it is going to "stay that way all the way thru life". Things change, like the port erosion conditions in carbines, and many people have just totally turned a deaf ear to that concern over the past years.
However, from an earlier post that I read here, it seems that overpressure ammo is being used in these guns, with pressures that are like 15k psi higher than they should be, and also higher than the specified max pressure rating of the bolt. Using 73k psi ammo, as alluded to in the earlier post, is a recipe for disaster. It's like using proof-loads for every shot. Not recommended practice. Anyone doing that has no room to be criticizing the bolt for failing.
Hmnn..... a bit large but informative
Failure Analysis of the M16 Rifle Bolt
V.Y. Yu*, J.G. Kohl, R.A. Crapanzano, M.W. Davies, A.G. Elam, M.K. Veach
Department of Civil and Mechanical Engineering
United States Military Academy
West Point, NY 10996, USA
Phone: 011.845.938.5507 Fax: 011.845.938.5522
Email: Victor.Yu@usma.edu*
Abstract
Recently, there have been several occurrences of failure in the bolt of the M16 rifle at a United States Army installation. Near the failure location, the bolt was subjected to repeated loading as the M16 was fired. In order to determine the stress distribution of the bolt due to the firing process, a geometric element analysis was performed using ProMechanica®. The fracture surface was examined using both an optical stereomicroscope and a scanning electron microscope in order to determine failure initiation and failure mode. It was discovered that the fracture initiated at a localized corrosion pit and propagated by fatigue. A controlled experiment, consisting of firing 1800 rounds using new bolts, showed that a region of wear developed near the site where fracture occurred in the failed bolt. This suggests that exposure of the base metal may have facilitated the formation of corrosion pits. In addition, Vickers microhardness profiles were taken on cross-sectional areas near the fillet region and 10 mm away from the failed locking lug. Disparities between microhardness profiles near the fillet region and 10 mm away from this region revealed that the bolt may not have been uniformly case hardened.
Keywords: failure analysis, abrasive wear, corrosion, geometric element analysis
1. Introduction
The M16 rifle was fielded into the U.S. Army in 1968 during the Vietnam War. The rifle has since been the primary assault rifle used by U.S. soldiers. The M16 has been through several modifications in its more than 30 years of service in the military. In July of 2003, an increasing trend in the amount M16A2 bolt failures was observed at a U.S. Army installation. Figure 1 displays the data of bolt failures observed over a five year span. These rifles were used over the past nine years during summer military training. As a result, this paper investigates the leading cause of catastrophic fracture of the bolt under firing conditions.
This study used both a geometric element analysis and a metallurgical analysis of the bolt. The goals of the methodology used are that 1) the geometric element analysis would reveal whether any elevated stresses existed in the bolt which would facilitate crack initiation and propagation; and 2) the metallurgical analysis would determine the fracture origin and failure mechanism. The metallurgical analysis would also determine whether mechanical properties of the material were insufficient for the designed operation of the bolt.
A controlled experiment was conducted which consisted of firing 1800 rounds using new bolts. After 1800 rounds, it was observed that there existed wear patterns which exposed new base metal to the environment at the same location as the failure initiation site on the fractured bolt. This exposed base metal may therefore serve as a site for corrosion pitting.
2. Geometric Element Analysis
2a. Procedure. In order to analyze the stresses that the bolt experienced while firing, a three-dimensional model of the bolt generated in Pro-Engineer® was used [1]. Figure 2 displays the three-dimensional model of the bolt. Subsequently, Pro-Mechanica® was used to post-process the model in order to calculate the von-Mises stresses in the bolt. Pro-Mechanica® differs from traditional finite element packages in that it does not use linear shape functions. Instead, Pro-Mechanica® fits polynomials up to 9th order as the shape function and is termed geometric element analysis. Thus, geometric element analysis offers accurate computational results even if the mesh is coarse, since the polynomials offer better convergence to the shape functions. As a result, the generated model of the bolt does not use purely linear shape functions since it can incorporate complex polynomials to fit the shape function. Furthermore, since a coarser mesh can be used to generate an accurate approximation of the model, bi-linear quadrilaterals were not solely used; instead, a mix of triangular elements and bi-linear quadrilaterals were incorporated into the model as shown in Figures 3a and 3b.
From historical data at the U.S. Army’s Testing and Armament Command (TACOM) and the Army Research Laboratory, a stress of 414 MPa was used to model the instantaneous force of the propellant combustion of the 5.56 mm round in the M16 rifle on the face of the M16 bolt. The conventional method for converting this type of dynamic process to a static analysis assumes that during the actual firing of the weapon, the pressure in the cartridge after combustion is dissipated by the rearward motion of the bolt. In order to conduct a static analysis of the bolt, half of the cartridge pressure was used to model this dissipation of energy [2]. Therefore, a stress of 207 MPa was used as a distributed load on the face of the bolt in the model. This force modeled the impact of the propellant igniting and exploding within the combustion chamber without incorporating the effects of recoil and the buffer assembly in the rifle. In addition, boundary constraints were placed on the bolt which allowed for minute deformations that the bolt would experience when the cartridge exploded in the combustion chamber.
2b. Results and Discussion. The von-Mises stress distribution in the bolt showed high stress concentrations present at the fillet of the locking lugs as shown in Figures 4a and 4b. In particular, higher stress concentrations were present in the locking lugs which were immediately adjacent to the round extractor. These two specific locking lugs experienced stresses on the magnitude of approximately 1070 MPa as shown in Figure 4b. All of the five fractured bolts analyzed at the Army installation had fractured at these specific locking lugs. Figure 5a shows a picture of a fractured bolt specimen and Figure 5b shows a picture of the fractured specimen at higher magnification. In addition, these extremely high stress concentrations contributed to the crack initiation which is evidenced by the picture of a crack growing from the locking lug next to the round extractor, as shown in Figure 6.
3. Metallurgical Analysis
3a. Procedure. The M16 bolt was also analyzed from a metallurgical viewpoint. This analysis determined whether additional factors other than stress concentrations contributed to the bolt failure. A chemical analysis of the bolt was conducted to determine if the material specifications were met, as displayed in Table 1. A stereomicroscope and a SEM were used to locate the fracture origin and to evaluate the fracture surface.
Vickers microhardness indentation was performed on a cross-sectional area near the fillet of the lug and at approximately 10 mm away from the lug on the bolt. Indentation profiles, consisting of five indents for each location, were taken which started 0.5 mm from the surface of the bolt and proceeded inward every 0.5 mm. Hardness readings are shown in Table 2.
In addition, a controlled experiment was conducted where three new bolts were subjected to the firing of a total of 1800 rounds. The experiment entailed firing the bolts in 300 round increments and subsequently cleaned with Royco 634 cleaner, lubricant, and preservative (MIL-PRF-63460D AM6) after each iteration. After the 1800 rounds were fired, the surface of each bolt was then examined using a stereomicroscope to detect any surface anomalies which might have occurred.
3b. Results and Discussion. Chemical analysis of the bolt composition revealed no significant differences between the failed bolt and Carpenter Steel 158 specifications [3], as shown in Table 1. Micrographs from a SEM revealed that the M16 bolt experienced corrosion in the form of localized pitting, as shown in Figure 7, near the locking lugs adjacent to the round extractor. From the SEM micrographs, the circumference of the fracture surface of the ruptured locking lug possessed shear lips. The existence of the shear lips signified ductile failure at the surface. However, the region at the corrosion pit did not have this characteristic shear lip. The absence of the shear lip at this location indicates that the bolt material was discontinuous at the surface. This discontinuity suggests that the corrosion pit is where failure initiated. The corrosion pit provides an additional stress concentration which aids in the initiation of the crack. In addition, the SEM micrograph displayed the presence of chevrons as seen in Figure 7. The chevron markings point back to the localized pit which further confirmed that the pit was the site for crack initiation.
Near the initiation site, the fracture surface was transgranular with faint fatigue striations indicating fatigue crack growth, as shown in Figure 7 and 8. Approximately 2.5 mm from the crack initiation site, the fracture surface transitioned from a smooth surface to a dimpled surface. This dimpled surface signified that the crack experienced unstable crack growth, or ductile failure, in this region.
The Vickers microhardness indentations taken at both locations show that the hardness reading is higher at the surface than towards the center of the bolt. This demonstrates that the surface was case hardened. However, the Vickers microhardness at the surface near the lug’s fillet was 100 units less than the hardness readings 10 mm from the fillet region. This indicates that the bolt was not uniformly case hardened. Thus, the softer region near the locking lugs is more susceptible to wear.
After 1800 rounds were fired using the new bolts, wear was observed which exposed the Carpenter steel 158 base metal to the environment, as shown in Figure 9. This area of observed wear on the surface of the bolt was in the same location as the crack initiation site on the fractured bolt, namely in the fillet region of the locking lugs adjacent to the round extractor. The base metal exposed due to the wear makes this specific area highly susceptible to corrosion pitting.
4. Conclusions
The fracture of the M16 bolt resulted from a cumulative effect of high stress concentrations at the fillet radius and the additional stress concentration imposed by the presence of localized pitting at the surface. The bolt possesses many fillet regions which impose numerous areas of high stress concentration. In particular, two fillets experienced higher stress immediately adjacent to the round extractor due to the non-contiguous feature of the bolt. These two specific areas of high stress concentration also corresponded to the same location where failure of the bolt occurred in all fractured bolt specimens. Micrographs obtained from the scanning electron microscope of the fractured surface showed localized pitting at the failure initiation site. In addition, transgranular crack propagation near the pit formations in the fillet regions was observed. The localized pits formed near the locking lugs also served as high stress concentration points. The presence of pits in the material amplified the stresses of the bolt in the locking lug region which already had a high stress concentration due to the irregular geometry of the bolt. This cumulative stress concentration provides a good indicator why the crack initiated and propagated from this region.
The wear observed in the controlled experiment indicates the mechanism of why the corrosion pits formed near the locking lug fillet by exposing the Carpenter Steel 158 base metal to the environment. Vickers microhardness readings near the fillet region show that the bolt was not uniformly case hardened. Comparison of the microhardness readings near the fillet region and 10 mm from this region show a disparity of approximately 100 units. The softer, less carburized region near the fillet contributes to the formation of a wear area after firing just 1800 rounds.
5. Acknowledgements
The authors would like to thank Mr. Victor K. Champagne, Jr. and the Materials Analysis Group at the Army Research Laboratory in Aberdeen Proving Grounds, MD for helpful discussions and for performing SEM work.
6. References
[1] Three-dimensional Pro-Engineer® model of M16 bolt from U.S. Army Testing and Armament Command, Rock Island, IL.
[2] Individual Weapon Systems & 3-D Technical Data Development Team, U.S. Army Testing and Armament Command, Rock Island, IL.
[3] Alloy data Carpenter No. 158® Alloy, Carpenter Technology Corporation, 1981.
Element Material Specifications [3] Percent Composition of Fractured Bolt
Carbon 0.10 0.185
Silicon 0.30 0.22
Nickel 3.50 3.43
Manganese 0.50 0.45
Chromium 1.50 1.39
Iron Balance Balance
Table 1. Chemical analysis of M16 bolt.
Distance from Bolt Surface (mm) Vickers Microhardness Reading near Lug Fillet Vickers Microhardness Reading 10mm from Lug
0.5 585.0 738.7
1.0 582.9 781.8
1.5 578.8 626.6
2.0 536.3 572.7
2.5 543.7 560.8
Table 2. Vickers microhardness readings of fractured M16 bolt on the cross-sectional area near lug fillet and 10 mm from lug fillet.
Figure 1. Historical data of M16 bolt failures at a U.S. Army installation.
Figure 2. Model of M16 bolt generated in Pro-Engineer®.
Figure 3a. Geometric Figure 3b. Geometric element mesh of locking lugs
element mesh of entire modeled in Pro-Engineer®.
bolt modeled in Pro-
Engineer®.
Figure 4a. Model of von-Mises stress
distribution in M16 bolt (above stress values
in Pa).
Figure 4b. Model of von-Mises stress distribution in
fillet region of locking lugs adjacent to round
extractor (above stress values in Pa).
Figure 5a. Fractured bolt specimen.
Figure 5b. Close-up of fractured lug.
Figure 6. Crack initiation and propagation in the
bolt specimen.
Figure 7. SEM micrograph of fractured
bolt surface.
Figure 8. Fatigue striations near fracture
initiation site.
Figure 9. Micrograph of wear area at fillet of locking lug after firing 1800
rounds.
Member
This issue of corrosion in the barrel near the gas tube.
Can that me managed with proper cleaning or is it inevitable?
I have been under the assumption that if I clean the weapon properly I should be able to avoid these types of issues.
Senior Member
Erosion, not corrosion.
A product of hot gases cutting steel like a cutting torch.
Member
double post.
Dealer
That is correct.
C4
Posting Permissions
Reply With Quote
Bookmarks